


IC 



9258 



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INF' : 




Mine Power Systems 



fHf^ 





YEAMS 



Information Circular 9258 



Mine Power Systems 



By Lloyd A. Morley 



UNITED STATES DEPARTMENT OF THE INTERIOR 
Manuel Lujan, Jr., Secretary 

BUREAU OF MINES 
T S Ary, Director 




Library of Congress Cataloging in Publication Data: 



c <3 



o ' 



Morley, Lloyd A. 

Mine power systems. 

(Information circular: 9258) 

Includes bibliographies. 

Includes index. 

Supt. of Docs, no.: I 28.27:9258. 

1. Electricity in mining. I. Title. II. Series: Information circular (United States. 
Bureau of Mines); 9258. 

TN295.U4 [TN343] 622 s [622'.48] 87-600213 



For sale by the Superintendent of Documents, U.S. Government Printing Office 
Washington, DC 20402 



PREFACE 



The application of electricity to the mining industry is a distinctive area of both 
mining engineering and electrical engineering. The difficult environment, the dynamic 
power loads, the cyclic and mobile operation and stringent safety requirements that 
characterize mining, all place unique demands on the mine power system. No other 
industry makes such extensive use of portable extensible equipment or has such com- 
plex grounding problems. Mine power systems can range from relatively simple in- 
stallations for small surface mines to complex underground systems where the harsh 
environment of dust, humidity, and cramped spaces stretches the ingenuity and crea- 
tivity of the engineer to provide reliable service. 

At the present time there is no up-to-date engineering text available that deals 
specifically with mine power systems. This has created extensive difficulties for edu- 
cators, industry engineers, and regulatory agency personnel. The need for a suitable 
reference for students in mining engineering provided the main impetus for this book, 
since the technician-level material that was in existence proved unsuitable for teach- 
ing young engineers who have little practical experience. 

The objective in preparing this manuscript was to assemble a single engineering 
reference on mine electrical power systems that is as comprehensive as possible. Ear- 
lier drafts of this material have been used successfully to instruct university students 
in courses ranging from basic electrical engineering through power-system design. It 
is felt, however, that the usefulness of this material extends beyond that of a student 
text. While not intended to replace other electrical or mining references, this publica- 
tion is also an indexed, reasonably comprehensive reference handbook for industry 
engineers and training personnel, and a source of material for electrical engineers who 
wish to expand their education into industrial power-system applications. Obviously, 
there will be some omissions; to include all aspects of mine electrical systems in one 
volume would approach an impossibility, but an attempt has been made to collect 
together the most significant information, thereby providing the tools needed to con- 
tinue a knowledgeable involvement in mine electricity. 

This reference work is divided into three general content areas. Chapters 1 through 
5 contain information considered elementary, chapters 6 through 11 deal with power- 
system components, and chapters 12 through 17 contain specifics on mine power sys- 
tems. A person familiar with electrical principals can use the earlier chapters as re- 
view material, but all chapters contain material relevant to mining and discuss the 
necessary combinations of equipment and components that should be contained in the 
mine power system. Emphasis throughout is placed on coal mining systems, although 
much of the material pertains to all mining operations. Both surface and underground 
power systems are discussed, the latter in more detail since these are the more com- 
plex systems and encounter the most problems. 

This publication is a thoroughly upgraded and extensively revised edition of Bureau 
of Mines Open File Reports 178(l)-82 and 178(2)-82, prepared under Bureau contract 
J0155009 by The Pennsylvania State University. It contains new chapters, new illustra- 
tions, and example problems that were not included in the original report. 

The assembly of this material has been a major undertaking. Many industry, 
academic, and Government agency personnel helped to review and critique practically 
every stage of draft preparation. The original report version was made available to 
students taking the mine power-systems courses at The Pennsylvania State Univer- 
sity, and their involvement was critical input to manuscript preparation. 

The author is grateful to all the companies and individuals who contributed or 
cooperated in this effort; so much information could not have been gathered without 
their help. A special thanks is owed to the late Robert Stefanko. He originally perceived 
the need for this text and provided guidance and encouragement throughout the proj- 
ect that produced the original report version. Others deserving special mention are A. 
M. Christman, R. H. King, J. A. Kohler, G. W. Luxbacher, T. Novak, J. N. Tomlinson, 
F. C. Trutt and D. J. Tylavsky. Each contributed directly to the text while on the fac- 
ulty or staff at The Pennsylvania State University; acknowledgements for their con- 
tributions are made in the individual chapters. 



11 



The author is grateful to the several individuals and companies that supplied 
noncopyrighted material for use in this publication. This material is noted by the vari- 
ous courtesies given throughout the text. Its incorporation does not constitute an en- 
dorsement by the author, The Pennsylvania State University, the University of Ala- 
bama, or the Bureau of Mines. 



Reference to specific products, equipment, or manufacturers does not imply endorsement by the Bureau of Mines. 



CONTENTS 



in 



Page 



Page 



Preface i 

Abstract 1 

Part I: Fundamentals 

Chapter 1.— Electrical power in mining 2 

Mine electrical history 2 

Underground mine history 2 

Surface mine history 4 

Mine power equipment 4 

Substations 5 

Switchhouses 5 

Power centers 5 

Distribution equipment 5 

Basic distribution arrangements 5 

Radial system 5 

Primary-selective system 6 

Primary-loop system 6 

Secondary-selective system 6 

Secondary-spot network 7 

Utility company power 7 

Surface mining 8 

Power systems in surface mines 8 

Main substations and subtransmission ... 8 

Surface mine distribution 9 

Underground coal mining 11 

Room-and-pillar mining 11 

Longwall mining 12 

Power systems in underground mines 13 

Regulations 13 

Underground mine distribution 13 

Surface facility power requirements 17 

Basic design considerations 17 

References 19 

Chapter 2.— Electrical fundamentals I 20 

Basic electrical phenomena 20 

Coulomb's law 20 

Voltage and current 20 

System of units 21 

Experimental laws and parameters 21 

Ohm's law 21 

KirchhofPs voltage law 22 

Kirchoffs current law 23 

Series circuits 24 

Parallel circuits 25 

The magnetic field 26 

Inductance 26 

Capacitance 28 

Electric field 28 



Instantaneous power 29 

Idealization and concentration 29 

Direct current circuits 30 

Direct current and circuit elements 30 

Series and parallel resistance 30 

Wye-delta transformations 33 

Circuit and loop equations 36 

Node equations 38 

Network theorems 40 

Time-varying voltages and currents 45 

Steady alternating current 48 

Effective alternating current 50 

Phasors 51 

Phasors and complex quantities 52 

Impedance transforms 53 

Steady-state analysis 55 

Chapter 3.— Electrical fundamentals II 59 

Average power and power factor 59 

Complex and apparent power 59 

Resonance 63 

Series resonance 63 

Parallel resonance 64 

Transformers 64 

Ideal transformer 66 

Actual transformers 68 

Conductor loss 68 

Leakage reactance 68 

Core losses and exciting current 69 

Power-transformer construction 70 

Transformer models 71 

Determination of transformer 

parameters 72 

Transformer efficiency and regulation ... 73 

Autotransformers 74 

Multivoltage transformers 74 

Current and potential transformers 75 

Chapter 4.— Power-system concepts 76 

Basic power circuit 76 

Three-phase circuits 76 

Balanced three-phase circuits 76 

Three-phase system voltages 77 

Load connections 78 

Line and phase currents 79 

Equivalent delta and wye loads 80 

Three-phase power 80 

Three-phase transformers 82 

Balanced three-phase circuit analysis 83 

One-line and three-line diagrams 85 



IV 



Page 



Page 



Circuits containing transformers 90 

Per-unit system 93 

Transformer impedance 94 

Three-winding transformers 95 

Per-unit method in system analysis 95 

Unbalanced three-phase circuits 97 

Fault types 98 

Fault analysis 98 

Symmetrical components 98 

Sequence components 99 

Sequence-quantity combinations 99 

Symmetrical-component relationship .... 100 

Symmetrical-component impedance 101 

Fault calculations 101 

Power terminology 102 

References 103 

Chapter 5.— Basic solid-state devices and 

instrumentation 104 

Semiconductors 104 

Diodes and rectifiers 104 

Diode equations 105 

Rectifier circuits 105 

Cooling 106 

Overloads 107 

Three-phase rectification 107 

Rectifier circuits 108 

Parallel rectifier operation 109 

Transistors 109 

Transistor operation 109 

Bipolar-transistor amplifiers 110 

Field-effect transistors 112 

Silicon-controlled rectifiers 113 

Integrated circuits 114 

Basic instrumentation 114 

Basic meter movements 115 

Meter-movement applications 116 

Wattmeters 117 

Varmeters 118 

Power-factor meters 118 

Power-system instrumentation 118 

Instrument transformers 118 

Single-phase connections 119 

Three-phase connections 120 

Special instruments 122 

Watthour meters 122 

Demand meters 122 

Bridges 122 

Megohmmeters 123 

Phase-sequence indicators 124 

Recording instruments 124 

Electronic instruments 125 



Electronic meters 125 

Oscilloscopes 125 

Tape recorders 126 

Transducers 126 

Instrument installations 127 

Part II: Power-System Components 

Chapter 6.— Motors and motor control 129 

Alternating current generation 129 

Principle of generator operation 129 

Generator construction 129 

Three-phase generation 131 

Direct current generators 131 

Motor basics 133 

Torque 133 

Speed-torque relationships 133 

Standardization 134 

Motor type 135 

Three-phase squirrel-cage induction 

motors 136 

Elementary three-phase motor 136 

Motor construction 138 

Motor behavior 138 

Insulation 139 

Design characteristics 139 

Induction-motor starting 141 

Wound-rotor induction motors 142 

Three-phase synchronous motors 143 

Synchronous-motor starting 144 

Synchronous-motor torque 145 

Generated voltage 146 

Power factor 146 

Applications 147 

Direct current motors 147 

Elementary motor 147 

Actual motor construction 148 

Torque 148 

Motor connections and performance .... 148 

Ward-Leonard system 152 

Mine motors 153 

Applications 153 

Actual equipment operation 153 

Single-phase motors 156 

Rotating stator field 156 

Split-phase starting 157 

Capacitor-start motors 157 

References 158 

Chapter 7. — Grounding 159 

Grounding systems 160 

Ungrounded neutral 160 



Page 

Solidly grounded neutral 160 

Low-resistance grounded neutral 160 

High-resistance grounded neutral 160 

Electric shock 161 

Characteristics of mine grounding systems . . 162 

Ground beds 162 

Grounding in underground mining 164 

Grounding in surface mines 166 

Ground-bed construction 166 

Ground resistance 166 

Electrode configuration formulas 167 

Two-layer earth structures 170 

Soil-heating effects 170 

Control of potential gradients 171 

Ground-bed resistance measurement 172 

Measurement method 172 

Ground test instruments 173 

Ground-bed resistivity 174 

Factors affecting resistivity 174 

Resistivity measurements 175 

Effect of chemical treatment of soils .... 177 

Ground-bed corrosion 177 

General ground-bed guidelines 178 

Grounding equipment 179 

Grounding resistor 179 

Grounding transformers 179 

Summary 180 

References 180 

Chapter 8.— Distribution 182 

Nature of cable distribution 182 

Cable components 183 

Conductors 184 

Insulation 185 

Cable jacket 185 

Cable shielding 186 

Cable types 186 

Cable terminations 191 

Cable couplers 191 

Coupler contacts 192 

Coupler insulation 192 

Coupler housing 192 

High-voltage couplers 193 

Low-voltage couplers 194 

Cable selection 194 

Cable length 195 

Conductor selection 195 

Cable installation and handling 202 

Borehole cables 203 

Feeder cable installation 204 

Recommended handling practices 204 

Cable failures and repairs 206 



Page 

Cable testing 206 

Failure location 207 

Splicing 207 

Trolley systems 211 

Trolley wire 211 

Trolley feeder 211 

Supports, lubrications, and turnouts 211 

Rails and bonds 215 

Overhead lines 216 

Overhead-line design 217 

Overhead-line electrocutions 218 

References 222 

Chapter 9.— Protective equipment and 

relaying 224 

Switching apparatus 224 

Arcs and circuit interruption 225 

Switches 226 

Circuit breakers 226 

Circuit breakers for low and medium 

voltage 227 

Molded case circuit breakers 228 

Power circuit breakers 232 

High-voltage circuit breakers 232 

Typical ratings 232 

Oil circuit breakers 232 

Minimum-oil circuit breakers 233 

Vacuum circuit breakers 234 

Fuses 235 

Low-voltage fuses 235 

Non-time-delay fuses 236 

Time-delay fuses 236 

Dual-element fuse 236 

Current-limiting fuses 236 

Standard fuses 236 

Nonstandard fuses 237 

High-voltage fuses 237 

Expulsion types 237 

Current-limiting high-voltage fuses 238 

Load-break switches 239 

Relays 240 

Relay terminology and types 240 

Thermal relays 240 

Electromagnetic-attraction relays 241 

Electromagnetic-induction relays 242 

Basic relay connections 244 

Alternating current direct relaying 244 

Alternating current potential relaying .... 246 

Alternating current differential relaying . . 247 

Direct current connections 247 

Kinds of protection 248 

Control wiring 248 



VI 



Page 

Phase protection 248 

Ground overcurrent 249 

Ground-check monitoring 251 

Advantages and disadvantages 253 

Arrangements for mining 254 

Zones of protection 254 

Coordination 254 

Ground-fault protection 254 

Overloads and short circuits 255 

Surface mines 255 

Underground mines 256 

References 259 

Chapter 10.— Sizing protective devices 260 

Fault current 260 

Fault-current sources 260 

Source equivalent circuit 260 

Fault calculations for three-phase systems . . 261 

Short-circuit calculation procedures 261 

Three-phase calculation example 264 

Computer fault analysis 268 

Ground-fault current calculations 268 

Direct current system faults 269 

Device settings 270 

Relay pickup settings 270 

Short-circuit protection 270 

Overload protection 271 

Ground-fault protection 271 

Current transformer matching 272 

Current transformer accuracy 272 

Accuracy calculations 273 

Low-voltage circuit breaker trips 274 

Overload protection 274 

Short-circuit protection 275 

Low-voltage power circuit breakers 275 

Fuses 276 

Coordination 276 

References 278 

Chapter 11.— Transients and overvoltages 280 

Transient sources 280 

Lightning phenomena 280 

Switching transients 281 

Capacitance switching 282 

Current chopping 284 

Prestrike 285 

Direct current interruption 286 

General switching transients 287 

Other transient phenomena 287 

Traveling waves 287 

Electromagnetic phenomena 290 

Transient-induced failures 290 



Page 

Winding response 290 

Coupling through transformers 291 

Transient protection 292 

Surge arresters 292 

Surge arrester applications 293 

Capacitors and system capacitance 295 

Other suppression devices 298 

Faraday shields 298 

Circuit arrangements 298 

Protection of overhead lines 298 

Impulse performance of ground beds .... 300 

References 301 

Part III: Mine Power Systems 

Chapter 12.— Mine power centers 302 

Equipment specifications 302 

Mine power centers 303 

High-voltage cable coupler 304 

Interlock switches 305 

Disconnect switch 305 

High-voltage fuses 306 

Surge arrestors 306 

Transformers 307 

Specifications 307 

Transformer construction 311 

Faraday shields 311 

Grounding resistor 311 

Busway 312 

Outgoing circuit breaker 312 

Ground-fault protection 314 

Single-phase transformers 316 

Metering circuits 316 

Outgoing cable couplers 317 

Ground-check monitors 317 

Power-factor correction 319 

Direct current utilization 320 

Rectifier transformer 321 

Rectifier 322 

Direct current ground-fault protection 

schemes 323 

Direct current control circuitry 324 

Direct current interrupting devices 324 

References 325 

Chapter 13.— Switchhouses and substations . . . 326 

Switchhouses 326 

Switchhouse internal components 326 

Switchhouse protective relaying 328 

Power circuit breakers 329 

Switchhouse control circuits 329 

Switchhouse design 331 



vu 



Page 

Substations 332 

Basic substation arrangements 332 

Single-ended substations 333 

Double-ended substations 334 

Substation transformers 334 

Substation switching apparatus 335 

Reclosers 335 

Disconnect switches and fuses 336 

Protective relaying in substations 336 

Lightning and surge protection in 

substations 337 

Substation grounding 338 

Substation ground mat 339 

Ground-fault protection 340 

Additional mine substation loads 340 

Portable substations 342 

Utility voltage as mine distribution 343 

Additional substation design 

considerations 344 

References 345 



Chapter 14.— Solid-state control and relaying 
Motor control 

Simple motor control 

Control systems 

Physical characteristics of thyristors . . . 

Direct current applications 

Alternating current applications 

Static protective relaying 

Operation of simplified solid-state and 

hybrid relays 

Static and electromechanical relay 

comparison 

Static relay mining applications 

Sensitive earth-leakage system 

Phase-sensitive short-circuit protection . 

Solid-state relays in the future 

Summary 

References 



346 
346 
348 
349 
349 
350 
351 
356 

356 

359 
361 
362 
363 
364 
364 
365 



Chapter 15.— Batteries and battery charging . . . 367 

Basic battery and battery-charging theory . . . 367 

Battery maintenance 370 

Chargers 370 

Charging stations 372 

Battery-box ventilation 374 



Page 

Battery surface leakage and faults 375 

Battery-charging hazards 377 

References 381 

Chapter 16.— Permissibility and hazard 

reduction 382 

Terminology 382 

Hazard-reduction methods 383 

Explosion-proof enclosures 383 

Explosion transmission 384 

Enclosure joints 385 

Enclosure mechanical strength and internal 

pressures 388 

Enclosure hazards 389 

Permissible equipment 391 

Permissible equipment schedule 391 

Maintenance of permissible equipment . . . 392 

Coal dust hazards 393 

Classifications of dust locations 393 

Reducing dust hazards 394 

Hazardous locations in preparation 

plants 394 

References 395 

Chapter 17.— Maintenance 396 

Mine maintenance program 397 

Economic justification 397 

Preventive maintenance program 

implementation 397 

Techniques of preventive maintenance 398 

Basic electrical measurements 398 

Insulation measurements 398 

Megohmmeter tests 400 

Mechanical measurements 404 

Continuous-monitoring systems 406 

Corona 406 

Corona behavior 408 

Corona detection 409 

Partial-discharge problems in mining .... 410 

Intermachine arcing 411 

Ground direct current offsets 412 

Summary 413 

References 414 

Bibliography 415 

Appendix.— Abbreviations and symbols 416 

Index 420 



ILLUSTRATIONS 



1.1. Simple mine electrical system arrangement 

1.2. Simple radial distribution system 

1.3. Power-center type of radial distribution . . 



Mil 



ILLUSTRATIONS-Continued 

Page 

1.4. Primary-selective distribution system 6 

1.5. Primary-loop distribution 6 

1.6. Secondary-selective system 7 

1.7. Secondary-spot network technique 7 

1.8. Representative utility transmission and distribution 7 

1.9. Subtransmission for surface mine 8 

1.10. Radial strip mine distribution system 9 

1.11. Secondary-selective distribution in strip mining 9 

1.12. Primary-loop design for strip mining 9 

1.13. Radial distribution for strip mine with overhead poleline base line 10 

1.14. Radial distribution for strip mine with all-cable distribution 10 

1.15. Surface mine distribution system using two base lines 10 

1.16. Open pit power system 11 

1.17. Layout of underground coal mine 11 

1.18. Plan view of retreating longwall 12 

1.19. Subtransmission for underground mine 13 

1.20. Radially distributed underground power system 14 

1.21. Secondary-selective distribution in underground mines 15 

1.22. Utilization in continuous mining section 15 

1.23. Power-system segment with longwall equipment 16 

1.24. Diagram of electrical-system segment for longwall 16 

1.25. Parallel-feed haulage system 17 

1.26. Representative expanded radial distribution for preparation plant 18 

1.27. Representative secondary-selective distribution for preparation plant 18 

2.1. Circuit element illustrating voltage polarity and current flow direction 22 

2.2. Simple series circuit 22 

2.3. Ideal and actual voltage sources 23 

2.4. Circuit for example 2.1 23 

2.5. Demonstration of Kirchhoff s current law 23 

2.6. Simple parallel circuits 24 

2.7. Ideal and actual current sources 24 

2.8. Parallel circuit for example 2.2 24 

2.9. Simple series circuit and equivalent 24 

2.10. Simple parallel circuit 25 

2.11. Series-parallel circuit for example 2.3 25 

2.12. Series-parallel circuit for example 2.4 26 

2.13. Magnetic flux in a straight conductor and in a long coil 26 

2.14. Demonstration of induced current 26 

2.15. Two coils demonstrating mutual inductance 27 

2.16. Long-coil inductance and inductor symbols 27 

2.17. Toroidal coil 28 

2.18. Charge, voltage, and current relationships of capacitor 28 

2.19. Electric lines of force between two parallel charged plates 28 

2.20. Resistor used to demonstrate instantaneous power 29 

2.21. Simple example of idealization and concentration 30 

2.22. Modeling of load center, trailing cable, and shuttle car 30 

2.23. Basic elements of resistance, inductance, and capacitance 31 

2.24. Simplification of dc circuit 31 

2.25. Simple circuit reduction 31 



IX 

ILLUSTRATIONS-Continued 

Page 

2.26. Circuit for example 2.5 32 

2.27. Circuit for example 2.6 32 

2.28. Series-parallel conductances for example 2.7 32 

2.29. Series-parallel circuit for example 2.8 33 

2.30. Two-terminal and three-terminal networks 33 

2.31. Wye and delta circuit configuration 34 

2.32. "T" and "tt" circuit configurations 34 

2.33. Common bridge circuit 34 

2.34. Circuit reduction of bridge circuit 35 

2.35. Parts of circuit 36 

2.36. Circuit demonstrating two independent loops 36 

2.37. Two-loop circuit for example 2.11 37 

2.38. Bridge circuit demonstrating loop analysis 37 

2.39. Three-loop circuit for example 2.12 38 

2.40. Simple two-node circuit 39 

2.41. Three-junction circuit 39 

2.42. Three-junction circuit with grounds 39 

2.43. Voltage-source circuit demonstrating node analysis 39 

2.44. Circuit for examples 2.13, 2.15, and 2.16 39 

2.45. Circuit for example 2.14 40 

2.46. Circuit for demonstrating superposition theorem 41 

2.47. Circuit in figure 2.44 with sources turned off 41 

2.48. Demonstration of reciprocity theorem 42 

2.49. Practical voltage-source model 42 

2.50. Practical current-source model 42 

2.51. Source transformation 43 

2.52. Circuit in figure 2.44 with current sources transformed to voltage sources 43 

2.53. Thevenin's theorem 43 

2.54. Norton's theorem 44 

2.55. Comparison of Thevenin's and Norton's circuits 44 

2.56. Circuit for example 2.17 44 

2.57. Active circuit for example 2.18 45 

2.58. Circuits illustrating solution steps to example 2.18 45 

2.59. Some time-varying electrical waves 46 

2.60. Sinusoidal ac waveform 46 

2.61. Steady ac showing phase shift 46 

2.62. Steady ac through resistance 46 

2.63. Steady ac through inductance 47 

2.64. Steady ac through capacitance 47 

2.65. Simple series RL circuit 48 

2.66. Simple series RC circuit 48 

2.67. Simple series RLC circuit 48 

2.68. Graphical representation of complex number 49 

2.69. Trigonometric or polar representation of complex number 49 

2.70. Sinusoid versus time and as phasor 51 

2.71. Phasor representation of current and voltage 51 

2.72. Other expressions for phasors 51 

2.73. Voltage-current phasor relationships for circuit elements 53 

2.74. Steady sinusoid analysis of simple RL series circuit 53 

2.75. Steady sinusoid analysis of simple RC series circuit 54 



ILLUSTRATIONS-Continued 

Page 

2.76. Steady sinusoid analysis of simple RLC series circuit 54 

2.77. Circuit for example 2.21 56 

2.78. Circuit for example 2.22 56 

2.79. Two-loop circuit for example 2.23 57 

2.80. Active circuit for example 2.24 57 

3.1. Power represented as real and imaginary components 60 

3.2. Illustration of leading and lagging power factors 61 

3.3. Circuit demonstrating sum of complex powers 62 

3.4. Simple series RLC circuit for resonance 63 

3.5. Plot of impedance magnitude versus frequency for series RLC illustrating resonance 63 

3.6. Circuits that exhibit parallel resonance 64 

3.7. Magnetic coupling between two conductors 64 

3.8. Magnetic coupling between two coils 65 

3.9. Demonstration of coil winding sense 65 

3.10. Dot convention for mutal inductance sign 65 

3.11. Demonstration of impedance transfer in transformers 67 

3.12. Ideal transformer with winding resistance included 68 

3.13. Accounting for transformer leakage flux 69 

3.14. Transformer magnetizing current 70 

3.15. Eddy current and magnetic hysteresis creating power loss in core 70 

3.16. Equivalent circuit of practical transformer 70 

3.17. Common power-transformer construction techniques 71 

3.18. Movement of exciting components to input 71 

3.19. Transferring secondary components to primary 71 

3.20. Final simplification of pratical circuit model 71 

3.21. Transformer parameter test series 72 

3.22. Circuit for example 3.8 73 

3.23. Comparison of two-winding transformer and autotransformer 74 

3.24. Two-winding transformer as an autotransformer 74 

3.25. Examples of transformers for multivoltage applications 75 

3.26. Two types of CT's 75 

3.27. Examples of CT and PT placement in circuit 75 

4.1. Basic power circuit 76 

4.2. Applications of basic power circuit 76 

4.3. Elementary three-phase generation 77 

4.4. Three-phase voltage sources 77 

4.5. Wye-connected source demonstrating line-to-line and line-to-neutral voltages 77 

4.6. Balanced three-phase load connections 78 

4.7. Four-wire wye-to-delta system 79 

4.8. Balanced delta load illustrating phase and line currents 79 

4.9. Comparison of equivalent delta and wye loads 80 

4.10. Three-single-phase transformers connected for three-phase operation 82 

4.11. Three-phase diagrams for the transformers of figure 4.10 82 

4.12. Open-delta three-phase transformer operation 83 

4.13. Per-phase reduction of wye-to-wye system 84 

4.14. Per-phase reduction of delta-to-delta system 84 

4.15. Three-line diagram 86 

4.16. One-line diagram of circuit shown in figure 4.15 86 



XI 

ILLUSTRATIONS-Continued 

Page 

4.17. Commonly used symbols for one-line electrical diagrams 87 

4.18. Symbols for relay functions 89 

4.19. One-line diagram for example 4.7 91 

4.20. Three-phase diagram of figure 4.19 91 

4.21. Per-phase diagram of figure 4.19 91 

4.22. One-line diagram with delta-delta transformer 92 

4.23. Per-phase diagram of figure 4.22 92 

4.24. One-line diagram with delta-wye transformer 92 

4.25. One leg of three-phase transformer from figure 4.24 92 

4.26. Approximate per-phase equivalent circuit for 750-kVA load-center transformer; impedance referred 

to high side 94 

4.27. Transformer of figure 4.26 with impedance referred to low side 94 

4.28. Simplified equivalent circuit of transformer expressed in per-unit 94 

4.29. Approximate equivalent circuit of three-winding transformer expressed in per-unit 95 

4.30. One-line diagram of small mine power system 95 

4.31. Impedance diagram of system in figure 4.30, expressed in per-unit on a 1,000-kVA base 97 

4.32. Basic fault descriptions 98 

4.33. Positive-sequence, negative-sequence, and zero-sequence vector sets 99 

4.34. Symmetrical component addition to obtain unbalanced three-phase set 100 

4.35. Equivalent delta-connected and wye-connected loads 100 

4.36. Three-phase system with line-to-neutral fault 101 

5.1. Symbol and operation of a p-n junction device 104 

5.2. Bias conditions and current flow for a diode 104 

5.3. Diode or rectifier characteristic curve 105 

5.4. Half-wave rectifier circuit and waveforms 106 

5.5. Single-way full-wave rectifier waveforms 106 

5.6. Bridge rectifier circuit and waveforms 106 

5.7. Example of filtering a rectifier output 106 

5.8. Heat sink cooling 107 

5.9. Heat sink thermal relationships 107 

5.10. Three-phase half -wave rectifier circuit and output voltage waveform 108 

5.11. Three-phase full-wave rectifier circuit with input and output voltage waveforms 108 

5.12. Parallel operation of rectifiers using paralleling reactors 109 

5.13. An n-p-n junction transistor 109 

5.14. A p-n-p junction transistor 109 

5.15. Current relationships for p-n-p and n-p-n devices 110 

5.16. Common-base amplifiers 110 

5.17. Common-emitter amplifier Ill 

5.18. Common-emitter characteristic curves Ill 

5.19. Bias techniques for common-emitter amplifiers 112 

5.20. Common-collector amplifier arrangment 112 

5.21. Model and symbols for junction FET's 112 

5.22. Example of a junction-FET application 113 

5.23. Model and symbols for MOS-FET devices 113 

5.24. SCR model and symbol 113 

5.25. SCR equivalent model and circuit 113 

5.26. General characteristic curve for SCR 114 

5.27. Sketch of simple monolithic IC cross section 114 

5.28. Top view of an actual IC 114 



xu 

ILLUSTRATIONS-Continued 

Page 

5.29. Examples of symbols employed for IC's 115 

5.30. Permanent-magnet moving coil movements 116 

5.31. Shunting d'Arsonval meter for high-current tests 116 

5.32. D'Arsonval meter used to measure dc potentials 116 

5.33. External shunts used for high-current measurements 117 

5.34. Simple ohmmeter circuit 117 

5.35. Rectifier ammeter 117 

5.36. Dynamometer connected as wattmeter 117 

5.37. Power-factor movement 118 

5.38. Simple instrument-transformer connections 118 

5.39. Voltmeter, ammeter, and wattmeter arranged as single-phase system 119 

5.40. Use of transducers with standard d'Arsonval movements 119 

5.41. Three-phase wattmeter connections 120 

5.42. Two-wattmeter method 120 

5.43. Three-phase power measurement with transducer 120 

5.44. Balanced three-phase measurement of voltage, current, and average power 121 

5.45. Line current measurements with two or three CT's 121 

5.46. Line-to-line voltage measurements with three or two PT's 121 

5.47. Simplified sketch of watthour meter induction mechanism 122 

5.48. Wheatstone bridge circuits 122 

5.49. Kelvin double bridge 123 

5.50. Megohmmeter testing insulation resistance 123 

5.51. Internal components of megohmmeter 123 

5.52. Phase-sequence indicator 124 

5.53. Strip-chart recorder 124 

5.54. Input circuits on electronic voltmeter 125 

5.55. Digital display 126 

5.56. Cathode-ray tube 127 

5.57. Semiconductor illustrating Hall effect 127 

6.1. Production of voltage from magnetic field 129 

6.2. Demonstration of ac generation 130 

6.3. Cross section of machine with salient poles on stator and nonsalient poles on rotor 130 

6.4. Cross section of machine with nonsalient poles on stator and rotor 130 

6.5. Simplified sketch of electromechanical machine illustrating physical components 130 

6.6. Elementary four-pole, single-phase ac generator 131 

6.7. Elementary two-pole, three-phase generator 131 

6.8. Elementary four-pole, three-phase generator 131 

6.9. Demonstration of dc generation 131 

6.10. Dc generator with two armature windings at right angles 132 

6.11. Separately excited dc generator 132 

6.12. Series dc generator 132 

6.13. Shunt dc generator 132 

6.14. Compound dc generator 132 

6.15. Current-carrying conductor in a magnetic field 133 

6.16. General speed-torque motor characteristic 134 

6.17. Examples of three frame number dimensions 134 

6.18. Demonstration of induction-motor operation 136 

6.19. Elementary three-phase induction motor 136 

6.20. Squirrel-cage rotor winding 136 



XU1 

ILLUSTRATIONS-Continued 

Page 

6.21. Rotating magnetic field in elementary three-phase, two-pole induction motor 137 

6.22. Induced rotor potential by rotating flux 137 

6.23. Lapped windings of three-phase motor stator 138 

6.24. Characteristic curves of three-phase induction motor 138 

6.25. Typical torque-speed characteristic for general-purpose induction motor 138 

6.26. Phasor diagrams of rotor and stator flux density for induction motor 139 

6.27. Typical torque-speed characteristics for NEMA-design three-phase squirrel-cage motors 140 

6.28. Other rotor-conductor designs 140 

6.29. Across-the-line magnetic starter 141 

6.30. Starting methods for induction motors 142 

6.31. Schematic of wound-rotor induction motor showing external resistance controller 143 

6.32. Torque-speed characteristics for wound-rotor motor with stepped-resistance controller 143 

6.33. Simplified step starter using individually timed magnetic relays 143 

6.34. Sketch showing construction of salient-pole synchronous motor 143 

6.35. Simplified diagram of synchronous motor using generator for field excitation 144 

6.36. External solid-state supply used to provide field excitation 144 

6.37. Schematic of low-speed cylindrical-rotor synchronous motor 144 

6.38. Controller used to demonstrate general starting method for synchronous motor 145 

6.39. Typical torque-speed characteristic for synchronous motor with damper winding 145 

6.40. Effect of load on rotor position 145 

6.41. Equivalent per-phase circuit of a synchronous motor and phasor diagrams for underexcited and 

overexcited field winding 146 

6.42. V-curves for synchronous motor 146 

6.43. Plan view of typical mining shovel showing m-g set 147 

6.44. Elementary two-pole dc motor 147 

6.45. Elementary four-pole dc motor 147 

6.46. Cross-sectional sketch of dc motor showing interpole and compensating windings 148 

6.47. Interaction between armature and main-field flux to produce main-field distortion 148 

6.48. Four connections for dc motors 149 

6.49. Typical characteristics for shunt, series, and compound motors of equal horsepower and speed 

ratings 149 

6.50. Simplified dc motor schematics with starting resistances 149 

6.51. Faceplate manual starter 150 

6.52. Multiple-switch starting 150 

6.53. Drum-type starter 150 

6.54. Simplified diagram of dynamic braking applied to shunt motor 151 

6.55. Two-step resistance starting of series-wound motor 151 

6.56. Forward-reverse switching of series-wound motor 152 

6.57. Dynamic braking applied to series-wound motor 152 

6.58. One-step starting of compound-wound motor 152 

6.59. Basic Ward-Leonard system 153 

6.60. Typical characteristic curves for each motor in traction locomotive 155 

6.61. Stator field of two-pole, single-phase induction motor 156 

6.62. Rotor field of stationary two-pole, single-phase induction motor 156 

6.63. Phase relationships between stator and turning rotor 156 

6.64. Starting and running stator windings 157 

6.65. Centrifugal switch to remove starting winding 157 

6.66. Capacitor-start motor 157 

7.1. Illustration of electrical shock hazard 159 

7.2. Capacitance coupling in ungrounded system 160 



XIV 

ILLUSTRATIONS-Continued 

Page 

7.3. Solidly grounded system 160 

7.4. Resistance-grounded system 160 

7.5. Effect of frequency on let-go current for men 162 

7.6. Simplified one-line diagram of substation 163 

7.7. Step potentials near grounded structure 163 

7.8. Touch potentials near grounded structure 163 

7.9. Line-to-earth fault resulting in current flow through safety ground bed 163 

7.10. Lightning stroke to equipment causing current flow through safety ground bed 164 

7.11. Lightning stroke current through system ground bed causing elevation of safety ground bed 164 

7.12. One-line diagram of simplified mine power system 164 

7.13. Mixed ac-dc mine power system; dc load energized from trolley system 165 

7.14. System grounding with current-limiting resistors 165 

7.15. Diode grounding of machine frame 165 

7.16. Resistance of earth surrounding electrode 166 

7.17. Decrease in earth resistance as electrode penetrates deeper soil horizons 167 

7.18. Calculated values of resistance and conductance for 3/4-in rod driven to depth of 25 ft 167 

7.19. Calculated values of resistance and conductance for 3/4-in rod driven to depth of 100 ft 167 

7.20. Nomogram to provide resistance of driven rod 168 

7.21. Resistance of one ground rod, 3/4-in diameter 168 

7.22. Resistance of parallel rods when arranged in straight line or circle with spacing equal to rod 

length 168 

7.23. Variation of earth resistance as number of ground rods is increased for various spacings between 

rods 168 

7.24. Values of coefficient V. x as function of length-to-width ratio of area 169 

7.25. Values of coefficient k 2 as function of length-to-width ratio of area 169 

7.26. Influence of first-layer height of potentials 171 

7.27. Potential on ground surface due to rod 6 ft long and 1-in diameter buried vertically at various 

depths 172 

7.28. Potential on ground surface due to strips, 1 in by 0.1 in, of various lengths buried horizontally at 

depth of 2 ft 172 

7.29. Measuring resistance of grounding system 173 

7.30. Concentric earth shells around ground connection being tested and around current electrode 173 

7.31. Correct spacing of auxiliary electrodes to give true resistance within 2.0% 173 

7.32. Resistivity range of some rocks, minerals, and metals 174 

7.33. Variation in soil resistivity with moisture content 175 

7.34. Typical resistivity curves of solutions 175 

7.35. Diagram for four-electrode resistivity survey showing lines of current flow in two-layer earth 176 

7.36. Connections for Wenner four-terminal resistivity test using megohmmeter 176 

7.37. Typical curve of resistivity versus electrode separation 176 

7.38. Reduction in ground mat resistance by soil treatment 177 

7.39. Seasonal resistance variations attenuated by soil treatment 177 

7.40. Trench model of soil treatment 177 

7.41. Voltage gradients in earth during ground-fault conditions 178 

7.42. Delta secondary with zig-zag grounding 180 

7.43. Delta secondary with wye-delta grounding transformer 180 

8.1. Cable distribution in underground coal mines 182 

8.2. Cable distribution in surface coal mines 183 

8.3. Shield types 186 

8.4. Cross sections of round unshielded mining cables 188 



XV 



ILLUSTRATIONS-Continued 

Page 

8.5. Cross sections of flat unshielded mining cables 188 

8.6. Cross sections of some shielded mining cables 188 

8.7. Round unshielded mining cables 189 

8.8. Flat unshielded mining cables 189 

8.9. Round shielded mining cables 189 

8.10. Cable types for typical distribution systems in underground coal mines 190 

8.11. Cable types for typical distribution systems in surface coal mines 190 

8.12. Cable terminations for applications up to 15 kV 191 

8.13. Coupler components 193 

8.14. Simplified one-line diagram for situation described in example 8.4 201 

8.15. Allowable short-circuit currents for insulated copper conductors 202 

8.16. Representative end-suspension termination for borehole cable 203 

8.17. Messenger wire supports for mine power-feeder cable 205 

8.18. Splice layout using template for staggered connections 208 

8.19. Effective method for removing unwanted insulation 208 

8.20. Staggering splice connections 209 

8.21. Examples of popular connectors and connections used in splices 209 

8.22. Reinstating power conductors with soft rubber tape 210 

8.23. Typical taped splice in high-voltage shielded cable 211 

8.24. Trolley-wire cross sections 212 

8.25. Typical trolley-wire and feeder-cable supports 214 

8.26. Trolley-wire semicatenary suspension 214 

8.27. Trolley system accessories 215 

8.28. Theoretical resistance of bonded joint 216 

8.29. Pole strength calculations 217 

8.30. Guy and log-anchor calculations 218 

8.31. Typical arrangements and pin-insulator spacings on wooded poles 218 

9.1. Typical system fault current 225 

9.2. Steps in circuit interruption 225 

9.3. Arc between two contacts 225 

9.4. Load-break switch 226 

9.5. Extinguishing arc by increasing the length 227 

9.6. Metal-barrier arc chute assists in arc deionization 227 

9.7. Insulated-barrier arc chute used with magnetic field 227 

9.8. Molded-case circuit breaker components 228 

9.9. Magnetic-trip relay 230 

9.10. Adjustable instantaneous setting 230 

9.11. Thermal-magnetic action of molded-case circuit breaker 230 

9.12. Time-current characteristics for thermal-magnetic circuit breakers 230 

9.13. Shunt-trip and undervoltage-release accessories 231 

9.14. Construction and operation of dead-tank OCB 233 

9.15. Turboaction are chamber for OCB's 233 

9.16. Cross section of minimum-oil breaker 234 

9.17. Cross section of VCB 234 

9.18. Operating mechanism for vacuum interrupter 235 

9.19. VCB assembly incorporating a load-break switch 235 

9.20. Common cartridge fuses 236 

9.21. Inside view of dual-element fuse 236 

9.22. Current-limiting action of fuses 237 



XVI 

ILLUSTRATIONS-Continued 

Page 

9.23. Energy-limiting action of fuses 237 

9.24. High-voltage power fuse and support 238 

9.25. Fusible element under spring tension in high-voltage fuse 238 

9.26. Cross section of boric acid power fuse refill 238 

9.27. Disassembled refill unit for boric acid fuse 238 

9.28. Load-break switch with interlocked high-voltage fuses 239 

9.29. Relay contact symbols 240 

9.30. Temperature-monitoring protector 240 

9.31. Electromechanical-thermal relays 240 

9.32. Solenoid and clapper relays 241 

9.33. Polar relay 242 

9.34. Common induction-disk relay 242 

9.35. Front view of induction-disk relay removed from case 242 

9.36. Inverse-time curve compared with definite-time curve 243 

9.37. Various time characteristics of induction units 243 

9.38. Family of inverse-time characteristics 244 

9.39. Cylinder directional relay 244 

9.40. Directional overcurrent relay using induction-disk relay and cylinder relay 245 

9.41. Direct relaying in ac system 245 

9.42. Potential-relaying connections 246 

9.43. Differential-relaying connections 247 

9.44. Dc direct-relaying connections 247 

9.45. Typical control wiring for UVR 248 

9.46. Typical control wiring for shunt-tripping element 248 

9.47. Three-phase overcurrent and short-circuit connections 248 

9.48. Two CT approaches 249 

9.49. Neutral-resistor current-relaying scheme 249 

9.50. Neutral-resistor potential-relaying scheme 250 

9.51. Zero-sequence ground relay connections 250 

9.52. Ground relay in residual connection 250 

9.53. Broken-delta protection 251 

9.54. Series loop ground-check monitor 251 

9.55. Transmitter loop ground-check monitor 252 

9.56. Bridge-type ground-check monitor 252 

9.57. Pilotless ground-check monitor 252 

9.58. Some difficulties associated with ground-check monitoring in mining 253 

9.59. Pilot interlocking circuit using ground-check monitor 254 

9.60. Simple surface mine power system illustrating protective relaying 255 

9.61. Typical schematic for three-phase molded-case circuit breaker with ground-overcurrent and 

ground-check protection 256 

9.62. One-line diagram of simple underground mine power system illustrating protective circuitry 257 

9.63. Diode-grounded system with possible fault indicated 257 

9.64. Basic grounding-conductor system 258 

9.65. Relayed grounding-conductor system 258 

9.66. Neutral-shift system 258 

9.67. Current-balance dc ground-fault relaying using saturable reactor 259 

9.68. Current-balance dc ground-fault relaying using saturable transformer 259 

10.1. Fault current waveform illustrating asymmetry 262 

10.2. Multiplying factors applied to three-phase faults to obtain momentary ratings for switching apparatus . . . 264 



xvu 

ILLUSTRATIONS-Continued 

Page 

10.3. Multiplying factors applied to three-phase faults to obtain close-and-latch ratings for switching 

apparatus 264 

10.4. One-line diagram for fault calculations 264 

10.5. Impedance diagram for one-line diagram of figure 10.4 266 

10.6. Simplification of figure 10.5 266 

10.7. Simplification of figure 10.6 267 

10.8. Further reduction of example network 267 

10.9. Equivalent circuit of figure 10.6 267 

10.10. Example problem with motor contribution neglected 267 

10.11. Network to calculate momentary or close-and-latch current duties 267 

10.12. Fault current in dc system 269 

10.13. Available fault current versus distance of fault from rectifier on typical trolley systems 269 

10.14. One-line diagram for pickup setting example 271 

10.15. Model of CT and its burden 272 

10.16. Typical set of saturation curves for 600/5 multiratio bushing-type CT 273 

10.17. Example of one-line diagram for preparing a coordination curve plot for one path 277 

10.18. Coordination curve plot for figure 10.17 showing various protective-device characteristics 277 

11.1. Schematic representation of lightning stroke discharge 280 

11.2. Distribution of crest currents in lightning strokes 281 

11.3. Map showing average number of thunderstorm days per year in United States 281 

11.4. Striking distances for negative and positive strokes 281 

11.5. Crest voltages induced on transmission lines by nearby strokes 281 

11.6. Simple circuit to illustrate capacitance-switching voltage transients 282 

11.7. Voltage and current waveforms before and after current interruption 282 

11.8. Voltage and current transient waveforms occurring with capacitance switching and restrike 283 

11.9. Per-phase diagram of 4,160-V pump-motor circuit 283 

11.10. Voltages and current wavesforms resulting from multiple restrikes after capacitance switching 284 

11.11. Graphic example of current chopping by breaker interruption 284 

11.12. Equivalent circuit of power-system segment with lumped components per phase, neglecting resistance . . 284 

11.13. Graphic example of chopping voltage transients 285 

11.14. Segment of mine power system 285 

11.15. Circuit to demonstrate voltage transients in dc system 286 

11.16. Transient overvoltage resulting from current interruption on dc system 286 

11.17. An undergrounded system, showing capacitive-current flow 287 

11.18. An undergrounded system, with fault on phase A 287 

11.19. The distributed inductance and capacitance of two-wire line shown as incremental sections 288 

11.20. Demonstration of traveling wave on overhead line 288 

11.21. Incident waves being reflected and refracted at discontinuity 289 

11.22. Electric field between conductors 290 

11.23. A 1.2 x 50 wave test used for BIL measurement 290 

11.24. Equivalent circuit of multiturn winding showing distribution inductance and capacitance 291 

11.25. Initial voltage distribution across uniform winding from step function 291 

11.26. Capacitive coupling of transient voltage through two-winding transformer 291 

11.27. Basic valve surge arrester 292 

11.28. Surge arrester with nonlinear resistance grading to equalize each gap structure 293 

11.29. Surge approaching surge- arrester-protected equipment 294 

11.30. Typical surge protection of rotating machinery and dry-insulated transformers 295 

11.31. Simplified sketch of mine power-system segment 296 



Will 

ILLUSTRATIONS-Continued 

Page 

11.32. Capacitance for 2,300-V induction motors 297 

11.33. Capacitance for 2,300-V synchronous motors 297 

11.34. Overhead ground-wire shielding for low and high distribution towers 299 

11.35. Static-wire-protection designs of wooded support structures using 30 protective angle 299 

11.36. Ratio of impulse to 60-Hz resistance as a function of peak impulse current, for driven rods 300 

11.37. Impulse breakdown of sand for two moisture conditions using spherical electrodes 300 

11.38. Impulse characteristics of spherical electrode, with seven attached pointed protrusions of various 

lengths 300 

12.1. Typical power centers used in underground coal mines 304 

12.2. Schematic illustrating major components in power center 304 

12.3. Top view of mine power center showing placement of many internal components 305 

12.4. Interconnections between input and feedthrough receptacles 305 

12.5. Graph illustrating transient crest voltage caused by ribbon-element current-limiting fuse operation 306 

12.6. Comparison of transformer withstand characteristic and surge arrester withstand characteristic 307 

12.7. Typical primary winding taps on power cable transformer 308 

12.8. Zig-zag grounding transformer 309 

12.9. Delta-wye connection for deriving a neutral , 309 

12.10. Technique for measuring transformer impedance 309 

12.11. Typical X/R ratio versus transformer capacity 310 

12.12. Typical mine power-center transformer under construction 311 

12.13. Completed transformer prior to installation 311 

12.14. Typical bus work in power center under construction 312 

12.15. Typical conductor connection to molded-case circuit breaker 313 

12.16. Zero-sequence relaying on outgoing circuit with control connections to breaker 314 

12.17. Zero-sequence relaying with jumper in relay case 314 

12.18. Neutral relaying applied to grounding-resistor current as backup protection 315 

12.19. Backup protection devices associated with mine power cables 315 

12.20. Typical test circuit for zero-sequence relaying 315 

12.21. Simple control circuit incorporating one ground-fault relay and one ground-check relay 316 

12.22. Simple convenience-outlet circuit for 120- or 240-V single phase 316 

12.23. Fuse mountings 316 

12.24. Typical metering circuit for line-to-line voltages 317 

12.25. Typical metering circuit for line currents 317 

12.26. Typical impedance monitor circuit 318 

12.27. Block diagram of continuity monitor connected in pilotless mode 318 

12.28. Block diagram of continuity monitor wired for pilot operation 318 

12.29. Application of power-factor correction in mine power center 320 

12.30. General arrangement of dc components for combination power center 320 

12.31. Full-wave bridge rectifier 321 

12.32. Series reactance to reduce available short-circuit current 321 

12.33. Separate transformer to increase impedance of dc circuit 321 

12.34. Typical full-wave bridge rectifier with two diodes in parallel per leg 322 

12.35. Diode with RC snubber protection 322 

12.36. Diode-grounded system 323 

12.37. Basic grounding-conductor system 323 

12.38. Relayed grounding-conductor system 323 

12.39. Neutral-shift system 323 

12.40. Differential current scheme 323 

12.41. Representative control circuit for rectifier 324 

12.42. Cross section of dc contactor 324 



XIX 

ILLUSTRATIONS-Continued 

Page 

13.1. Diagram for typical single switchhouse 326 

13.2. Control circuitry for single switchhouse using battery tripping 326 

13.3. Diagram for typical double switchhouse 327 

13.4. Control circuitry for double switchhouse using capacitor tripping 327 

13.5. Typical family of curves for inverse-time relay 328 

13.6. Illustration of fault location for adjusting selectivity 328 

13.7. Typical control circuit for double switchhouse using capacitor tripping 330 

13.8. Typical control circuit for single switchhouse using battery tripping 331 

13.9. Overall view of main substation serving mine 332 

13.10. Radial distribution applied to underground mine and its surface facilities 333 

13.11. One-line diagram for single-ended substation with fuse-protected transformer 333 

13.12. One-line diagram for single-ended substation with circuit-breaker-protected transformer 333 

13.13. Simplified one-line diagram for doubled-ended substation 334 

13.14. Typical liquid-immersed transformer in substation 334 

13.15. Dead-tank OCB in substation 335 

13.16. Standard percentage-differential relaying system for transformer protection 336 

13.17. One-line diagram of substation with percentage-differential relaying 337 

13.18. Insulation characteristic of liquid-immersed transformer compared with the characteristic of valve 

surge arrester 338 

13.19. Plan view showing locations of system and safety ground beds 338 

13.20. Typical system ground bed for large substation 340 

13.21. Typical system ground bed for small substation 340 

13.22. Substation feeding both surface and underground loads (no grounding conductor) 341 

13.23. Substation feeding both surface and underground loads 342 

13.24. Typical portable substation to service small mine 343 

13.25. Providing mine ground and protective relaying from utility substation 344 

13.26. Use of isolation transformer with utility substation 344 

14.1. Model and circuit symbol for thyristor 346 

14.2. Typical characteristics curve for thyristor 346 

14.3. Thyristor half-wave rectifier 347 

14.4. Alternating current thyristor control 347 

14.5. Three-phase control with bidirectional thyristor arrangement 348 

14.6. Full-wave thyristor bridge rectifier 348 

14.7. Three-phase thyristor-controlled rectifier 348 

14.8. Simplified chopper control 348 

14.9. Basic control-system block diagram 349 

14.10. Simplified block diagram of a motor controller 349 

14.11. Common thyristor configurations 349 

14.12. Heat sinking of disk-type thyristors 349 

14.13. Block diagram of ac-dc shuttle car 350 

14.14. Block diagram of ac-dc continuous miner 351 

14.15. Simple variable-frequency control 351 

14.16. Elementary inverter circuit 351 

14.17. Use of variable-frequency drive on production mining shovel 352 

14.18. Simplified diagram of current-regulated static belt starter 353 

14.19. Simplified diagram of linear-acceleration static belt starter 353 

14.20. Types of thyristor firing pulses 355 

14.21. Thyristor protection for static belt starters 355 

14.22. Protective-relay connections 356 

14.23. Simple electromechanical relay 357 



XX 

ILLUSTRATIONS-Continued 

Page 

14.24. Simple static relay 357 

14.25. Transistor used as relay : . . . 357 

14.26. Optical transistor as relay 357 

14.27. Thyristor used as relay 357 

14.28. Triac used as relay 358 

14.29. Hybrid static relays 359 

14.30. Simple overcurrent static relay 359 

14.31. Simplified sketch of the SEL system 362 

14.32. Simplified sketch of the multipoint SEL system 363 

14.33. Diode-bridge phase-sensitive protection 364 

14.34. Equivalent model of figure 14.33 364 

14.35. Electronic-comparator method of phase-sensitive protection 364 

14.36. Digital-controlled continuous static relay used for timed overcurrent 365 

15.1. Composition of lead-acid storage battery in various states of charge 368 

15.2. Voltage per cell of a typical lead-acid battery with varying continuous rates of discharge 369 

15.3. Typical charging process of cell from 18-cell, 725-Ah battery 369 

15.4. Simplified schematic of saturable-reactor charger 371 

15.5. Simplified schematic of single-phase thyristor charger 371 

15.6. Two-winding transformer model 371 

15.7. Representation transformer magnetization curve 371 

15.8. Ferroresonant transformer model 372 

15.9. Ferroresonant transformer 372 

15.10. Ferroresonant battery charger 372 

15.11. Plan of underground charging station 373 

15.12. Circuit for detecting faults in batteries 376 

15.13. Curve of relay current for various fault positions on battery 376 

15.14. One-line diagram of desired charger features 378 

16.1. Cross-sectional sketch of typical explosion-proof enclosure 384 

16.2. Typical plane-flange joint 385 

16.3. Typical step-flange joint 385 

16.4. Threaded joint 385 

16.5. Tongue-and-groove joint 385 

16.6. Blind screw hole 386 

16.7. Pressure vent limiting pressure buildup during internal explosion 387 

16.8. Pressure vent assembly using metal-foam material 387 

16.9. Typical slip-fit straight stuffing box and packaging-gland lead entrance 387 

16.10. Typical slip-fit angle stuffing box and packing-gland lead entrance with hose clamp 387 

16.11. Typical slip-fit angle stuffing box and packing-gland lead entrance 388 

16.12. Typical plug for spare lead-entrance hole 388 

16.13. Typical threaded straight stuffing box and packing-gland lead entrance with provision for hose 

conduit 388 

16.14. Prototype trailing cable entry with polyurethane grommet 388 

16.15. Insulated-stud lead entrance 388 

16.16. Decision flow chart of class II, division 1 and 2 hazardous locations 394 

17.1. Circuit modeling a dielectric 399 

17.2. Current-voltage characteristics in a dielectric 399 

17.3. Graph relating approximate insulation resistance variation with temperature for rotating machines 400 



XXI 

ILLUSTRATIONS-Continued 

Page 

17 A. Insulation resistance versus application time of test voltage 400 

17.5. Megohmmeter test connections for checking cable insulation in line A .... 401 

17.6. Megohmmeter test connections for ac motor 401 

17.7. Megohmmeter test connections for dc motor 401 

17.8. Spot resistance curve for normal motor 401 

17.9. Spot resistance curve showing effects of dust and moisture 401 

17.10. Spot resistance curve for detective motor 402 

17.11. Megohmmeter test connections for transformer 402 

17.12. Time-resistance curve 402 

17.13. Three time-resistance curves for deteriorating motor 402 

17.14. Time-resistance curves showing polarization for hypothetical motor 403 

17.15. Polarization factor curve for deteriorating motor 403 

17.16. Multiple voltage curves for deteriorating motor 403 

17.17. Circuit for harmonic tests 404 

17.18. Power-factor versus voltage curves showing tie-up 404 

17.19. Mounting techniques for two vibration transducers 405 

17.20. Four typical vibration measurement points 405 

17.21. Typical vibration severity chart 405 

17.22. Comparison of acoustic-emission techniques for detecting failing roller bearings 406 

17.23. Conceptual diagram of generalized mine monitoring and control system 406 

17.24. Conduction in gas 407 

17.25. Discharge sequence in an ionizing field 407 

17.26. High-stress geometries 408 

17.27. Typical dielectric voids in cables 409 

17.28. Block diagram for corona-detection system 409 

17.29. High-voltage cable terminations 410 

17.30. Major insulation void sometimes found in high-voltage coupler terminations 411 

17.31. Possible stress site in high-voltage coupler insulators 411 

17.32. Power-conductor transposition on three-conductor type G cable 412 

17.33. Application of diode-suppression bridges in power center 412 

17.34. Typical saturable-reactor characteristic 412 

TABLES 

2.1. SI symbols and units 21 

2.2. Resistivity of some common materials at 20 C 22 

4.1. IEEE device numbers and functions 90 

4.2. Device numbers and letters common to mining 90 

6.1. Motor voltage ratings common to mining 135 

6.2. Motor insulation classes 135 

6.3. NEMA class A standard starters for three-phase induction motors 141 

6.4. Common motors for mining equipment 153 

7.1. Current range and effect on a typical man weighing 150 lb 161 

7.2. Typical resistances for various contact situations 162 

7.3. Approximate resistance formulas for various electrode configurations 170 

7.4. Comparison of grounding grids with other types of electrodes 172 

7.5. General resistivity classification 174 

7.6. Variations in resistivity with geologic age 174 

7.7. Typical values of resistivity of some soils 174 

7.8. Variation in soil resistivity with moisture content 175 

7.9. Typical potentials of metals in soil measured from a copper and copper sulfate reference electrode .... 178 



XX11 

TABLES-Continued 

Page 

8.1. Conductor sizes and cross-sectional areas 184 

8.2. Letters used in alphabetic cable code 187 

8.3. Codes for typical cables used in mining 187 

8.4. Typical diameters for round portable power cables 193 

8.5. Typical diameters for flat portable cables 193 

8.6. Specifications for trailing cables longer than 500 ft 195 

8.7. Ampacities for portable power cables 196 

8.8. Ampacities for three-conductor mine power cables 196 

8.9. Correction factors for ampacities at various ambient temperatures 196 

8.10. Ampacity derating factors for 60 C-rated trailing cables operated on drums 197 

8.11. Australian specifications for ampacity derating factors for trailing cables operated on drums 197 

8.12. Some estimated power factors and load factors for various underground coal mining equipment in 

good operating conditions 198 

8.13. Intermittent-duty ratings for trailing cables 199 

8.14. Resistance and reactance of portable power cable 200 

8.15. Resistance and reactance of mine-power-feeder cable 201 

8.16. Solid-wire breaking strength 202 

8.17. Recommended minimum bending radius, unshielded or unarmored cables 204 

8.18. Recommended minimum bending radius, shielded and armored cables 204 

8.19. Trolley-wire specifications 212 

8.20. Characteristic data for solid copper feeder cable 213 

8.21. Characteristic data for stranded copper feeder cable 213 

8.22. Trolley-wire support spacings on curves 215 

8.23. Resistance of steel rail at 20 C 215 

8.24. Data for rail-bond cable 216 

8.25. Minimum vertical conductor clearances as specified by the NESC, applicable to mining and 

mining-related operations 220 

8.26. Minimum distances from overhead lines for equipment booms and masts 221 

9.1. Ratings for mining-service molded-case circuit breakers 228 

9.2. Interrupting-current ratings versus system voltage 229 

9.3. Maximum instantaneous-trip settings 230 

9.4. Commonly available magnetic-trip ranges for mining- service molded-case breakers 230 

9.5. Some typical ratings for low-voltage power circuit breakers 232 

9.6. Typical minimum-oil circuit breaker ratings 234 

9.7. Ratings of high-voltage power fuses 239 

9.8. Common current ratings of induction-disk overcurrent relays 243 

9.9. Standard burden for current transformers 246 

9.10. Standard ratings for potential transformers 261 

10.1. Sample reactances for synchronous and induction motors 261 

10.2. Three-phase transformer per-unit impedances for liquid-immersed transformers 262 

10.3. Three-phase transformers impedances for distribution transformers, including load centers 262 

10.4. Sample applications of fault calculations 263 

10.5. Impedance of cables in figure 10.4 265 

10.6. Burdens of relay elements and ammeter connected to CT's 273 

10.7. Recommended instantaneous trip settings for 480-, 600-, 1,040-V three-phase trailing-cable protection . . 275 

10.8. Recommended instantaneous trip settings for 300- and 600- Vdc trailing-cable protection 276 

11.1. Recommended station and intermediate surge arresters for resistance-grounded mine power systems to 

protect oil-immersed transformers 294 



XX111 

TABLES-Continued 

Page 

11.2. Recommended distribution-class, RM-type, surge arresters for resistance-grounded mine power systems 

to protect rotating machinery and dry-type transformers 294 

11.3. Commonly used surge capacitors for limiting voltage rate of rise on rotating machinery and 

dry-insulated transformers 295 

11.4. Typical capacitances per phase of power-system components, for shielded power cable SHD, SHD-GC, 

and SHD+GC 297 

11.5. Typical capacitances per phase of power-system components 297 

11.6. Protective angle versus structure height 299 

12.1. Typical current ratings of 400-A load-break switch 305 

12.2. Typical ratings for combination power centers 321 

13.1. Standard impedance for liquid-immersed three-phase transformers 335 

13.2. Standard BIL's for oil-immersed power transformers 337 

14.1. Typical electromechanical and static relay characteristics 358 

14.2. Time-margin comparison between electromechanical and static relays 360 

14.3. Comparison of induction-disk and static time-overcurrent relay burdens to a current transformer 361 

15.1. Formulas to estimate hydrogen evolution 373 

16.1. Structural gap dimensions for explosion-proof enclosures as specified by 30 CFR 18 385 

16.2. Minimum autoignition temperatures versus layer thickness for bituminous coals 393 

17.1. Common causes of vibration 405 



MINE POWER SYSTEMS 

By Lloyd A. Morley 1 



ABSTRACT 

This Bureau of Mines publication presents a comprehensive review of mine elec- 
trical power-system theory and practice. It discusses fundamental theory and the vital 
aspects to be considered in planning and designing mine electrical power systems. The 
report is divided into three major sections. The first presents the history of electricity 
in mining and the fundamentals of electrical phenomena and components. The second 
focuses on power-system components: motors, grounding systems, cables, and protec- 
tive equipment and devices. The final section includes mine power-center equipment, 
switchhouses and substations, batteries, and mine maintenance. 

'Professor of mining engineering, The Pennsylvania State University, University Park, PA (now professor and department head, mineral engineering, 
University of Alabama, Tuscaloosa, AL). 



CHAPTER 1 .—ELECTRICAL POWER IN MINING 



Probably no other mining area has grown so rapidly yet 
been as little understood by the average mine worker or 
operator as the mine electrical power system. Traditionally, 
the field has held little interest for the mining engineer, 
who has tended to avoid it, or for the electrical engineer, 
who has given it scant attention. But today's mine power 
system is both complex and subject to numerous legal con- 
straints, and it is no longer possible to treat it with the 
indifference of the past. 

Underground mining machines are among the most 
compact and rugged equipment over designed, and individ- 
ual units can have up to 1,000 total horsepower. Mining 
equipment is usually mobile and self-propelled; most is 
powered electrically through portable cables and, for safety, 
must be part of an elaborate grounding system. The ma- 
chines and power-distribution equipment are seldom sta- 
tionary, must be adapted to continuous cyclic operation, and 
must resist daunting levels of dust and vibration. 

Surface mining can involve the largest earth-moving 
equipment built, where one piece can have 12,000 or more 
connected horsepower— the largest today is over 30,000 hp. 
The electrical loads created by this machinery are cyclic 
and extremely dynamic: the largest excavator, for exam- 
ple, can require electrical loads that range from 200% 
motoring to 100% generating every 50 to 60 s, under the 
most exacting physical conditions. In the ever-moving min- 
ing operation where distribution of power must be con- 
stantly extended and relocated, subjected to abuse by 
machine and worker alike, the potential for safety hazards 
is always present. 

Engineering and maintaining such an electrical system 
is demanding and challenging. It requires a specialist with 
knowledge of both mining and electrical engineering. Yet 
conversely, the effective management of a mine requires 
that anyone responsible for production and safety also be 
conversant with the mine electrical system. Management 
should understand the advantages and disadvantages of one 
system over another, for if the power system is poorly de- 
signed, not only will safety be compromised but the mine 
operator will pay for the resulting conditions with high 
power bills, high-cost maintenance, and loss of production. 

Too often, a new mine is designed to use the type of 
power system employed in the preceding mine, without a 
comprehensive power study to determine the system needs 
and examine the alternatives available. Problems arise in 
existing mines when new mining equipment has been 
adopted without due regard for its impact on the operating 
power system; these problems haunt the mine electrical 
engineer who must frequently cope with a system that is 
a mongrel, bred from diverse inheritances from the past 
combined with recent changes and additions. New laws, 
standards, and safety requirements must frequently be 
accommodated by power systems not originally designed 
to meet their specifications; new and unfamiliar equipment 
must be grafted to the existing network, and the result can 
be a hybrid of considerable complexity. This text has been 
produced to assist the power engineer and the student in 
understanding these complexities and the principles that 
lie behind them. 

The material presented here is structured so an indi- 
vidual unfamiliar with electrical engineering can first 
develop the necessary fundamentals before embarking into 
mine electrical design. A basic physics and calculus knowl- 



edge is necessary to understand the content completely. The 
goal has been to assemble the most significant information 
required for comprehension of mine power systems so that 
the reader may then progress to more specialized topics. But 
first, a brief review of the development of electrical usage 
in mines is given, in order that the reasons for some of the 
peculiarities of mine power systems can be appreciated. 



MINE ELECTRICAL HISTORY 

Electricity was first introduced into coal mines shortly 
before the beginning of the 20th century in the form of di- 
rect current (dc) for rail haulage. This form of current was 
used because at that time most systems were powered by 
dc generators. It had a number of advantages for haulage; 
the most outstanding was that the dc series-wound motor 
had (and has) excellent traction characteristics. Speed con- 
trol was a simple matter of placing a resistance in series 
with the motor armature or field circuits. 

Batteries served as the first power source, and hence 
the vehicle was extremely mobile even though constrained 
on rails. However, keeping the batteries charged was both- 
ersome, so trolley wires were soon introduced in several 
mines. Allowing the trolley wire to act as one conductor and 
the rail as the other provided the simplest form of power 
distribution yet known to the mining industry. Available 
haulage machinery of that period was low in horsepower 
and the mines were relatively small so the increased resis- 
tance that reduced voltage and power supplied to the motors 
was still acceptable. Thus, the dc system at a voltage of 250 
or 550 V became firmly entrenched in coal mines. 

Underground Mine History 

Underground, the first electrically driven coal mining 
machine, the coal cutter, was installed in the early 1920's. 
Although dc offered no special advantage, it was readily 
available; hence, the machine was equipped with a dc motor 
and added to the system. The cutter was followed almost 
immediately by the loader, and it too was driven by dc 
motors. If there was rail haulage in the mine, trailing cables 
supplied power from the trolley wire and the rail to the 
machines. 

The next significant increase in power consumption 
came with the introduction of the shuttle car, almost 20 yr 
after the coal cutter. Actually, when the shuttle car was 
first invented in 1937, it was battery powered. The addi- 
tion of an automatic reeling device to handle a trailing cable 
came later, in an attempt to overcome battery deficiencies. 
These trailing cables were also connected to the haulage 
power system, and this equipment, when combined with the 
cutters and loaders, placed additional stress on the dc 
distribution system. 

At that time, the horsepower required to operate each 
piece of electrical mining equipment was quite small and 
no individual machine used a large amount of current. 
However, when all machines were combined, significant 
power was required, and because all the conductors offered 
resistance, voltage drops and transmission losses in the 
distribution system were extensive. Alternating current (ac) 
would have been more practical because it could have been 



distributed easily at a higher voltage, thereby reducing cur- 
rent, voltage drops, and transmission losses. But many 
States had stringent limitations on maximum voltage 
levels, usually around 300 V, and with this restriction ac 
had no advantage over dc. Hence, dc continued to be used 
to operate the successful combination of cutters, loaders, 
and shuttle cars. 

Development in ac-to-dc conversion equipment played 
an important role in underground coal mine power utiliza- 
tion throughout this period. Motor-generators or synchron- 
ous converters were originally employed for conversion pur- 
poses, but in addition to being heavy and bulky, they could 
not be operated effectively in a gassy and dusty atmosphere, 
and maintenance requirements were substantial. As a re- 
sult, most conversion installations were placed on the sur- 
face with borehole connections to the underground mine. 
This was acceptable as most mines were then relatively 
shallow. 

In the 1930's, the same decade that saw the inception 
of the shuttle car, mercury-arc-ignition rectifiers began to 
be employed to provide dc underground. The arc tubes al- 
lowed more efficient use of electricity in deeper and larger 
mines than had previously been possible. As the tubes had 
no moving parts, maintenance was lower, efficiency was 
higher, and portability was improved. These rectifiers were 
usually centrally located in the mines because a liquid heat 
exchanger made them heavy and bulky. In this way, dis- 
tribution to the mine rectifier was ac, but distribution 
throughout most of the mine electrical system was still dc. 
At about the same time, some mines found that haulage 
of materials by conveyor could be more efficient than haul- 
age by rail. The conveyors were also powered by dc motors, 
and stress continued to be added to the electrical system. 

In the late 1940's, when continuous mining machines 
first began to be used extensively, dc was again expected 
to provide the power. However, the continuous miners nor- 
mally needed more energy input than the sum of the various 
conventional mining equipment they replaced, and because 
the required horsepower created high current demand, dc 
was found to be entirely unsatisfactory in most cases. The 
attendant current demand created enormous voltage drops 
in the distribution system. As a possible solution, the dc 
supply system was separated from the haulage system, but 
eventhis was unable to improve voltage regulation. During 
peak operation periods, voltages at the machines were so 
far below the values called for that even moderate efficiency 
was impossible. The increasingly large cable sizes required 
to supply the needed power created difficult cable-handling 
problems. The use of three-phase ac distribution and motors 
was an obvious answer, but for at least a decade some min- 
ing companies were reluctant to make the change. In many 
instances this was because the laws in some States limited 
maximum voltages in the mine. Lawmakers were convinced 
that high voltages were synonymous with high safety risks. 
Some State laws were not updated until the mid 1960's. 

When higher voltages were finally permitted, the de- 
sirable economics of ac employment could be realized and 
there was a swift transformation from dc to ac for both 
distribution and high-horsepower loads in underground coal 
mines. Unfortunately, many mine electrical systems were 
at least partially modified without concern for the compat- 
iblity of these changes with the remainder of the system, 
and various safety and production problems arose. 

As a result of conversions, mine power systems gener- 
ally had two voltage levels, one for distribution and one for 
utilization. The simplified mine electrical arrangement 



shown in figure 1.1 illustrates the results. Here, the sub- 
station transforms the utility voltage down to distribution 
levels, which are most often at high voltage greater then 
1,000 V. Power at this voltage is transmitted or distributed 
through conductors from the substation to the power center; 
hence, this system is called the distribution system. The 
power center or load center, actually a portable substation, 
transforms the voltage to utilization levels, which are 
typically at low voltage of 660 V or less, or medium vol- 
tage of 661 to 1,000 V. At this level, or face voltage, power 
is normally delivered to the equipment. Despite this ref- 
erence to voltage levels, it should be noted that distribu- 
tion and utilization describe functions of a power system 
segment, not specific voltage ranges. 

Originally, primary ac distribution was made at 2,300 
or 4,160 V. In most mines, these levels were later increased 
to 7,200 V. Some operations recently increased the voltage 
to 12,470 or 13,200 V for both longwall and continuous- 
mining applications. Each new distribution voltage, it may 
be noted, is a factor of \pi times the previous value 
(\/ r 3(2,300) = 4,160). The principal reason for increasing the 
voltage was that, for the same load, current would be cor- 
respondingly smaller, and lower distribution losses would 
result even though the same conductor sizes were used. 

From the beginning, 440 Vac was the most popular 
voltage for utilization, despite the fact that when the con- 
tinous miner proved so successful its horsepower was pro- 
gressively increased, following the sometimes misguided 
notion that a directly proportional increase in coal produc- 
tion would follow. As with dc, the additional horsepower 
resulted in an increase in trailing-cable sizes, until the 
weight of the cables was almost more than personnel could 
handle. To compensate, the most common move was to raise 
the rated motor voltages to 550 Vac. More recently, manu- 
facturers have produced machines with 950-V (550 \/15) 
motors to further overcome the trailing-cable problems. 

While these changes were being made to ac for machine 
operation and distribution, the use of dc for haulage con- 
tinued to be advantageous. In the early 1960's, silicon diode 
rectifiers with large current capabilities became available. 
Simple, efficient, and small, these rectifiers were ideally 
suited for use underground and made ac distribution possi- 
ble for the entire electrical system except rail haulage. 
Through the use of rectifiers, the benefits of dc for traction 
and of ac for distribution and utilization on high power loads 
could be realized. For example, while continuous miners 
normally used ac, part of the supply at the power center 
was rectified to dc, primarily for powering the shuttle cars. 

These underground electrical systems appeared to be 
simple, and as a result they did not become the focus of at- 
tention for some time. Systems were frequently designed 
and maintained by a "seat-of-the-pants" approach, to the 
point that ac distribution and equipment were installed 



source Substation Switchhouse 



Power 
center 



o- 



Distribution 
voltage 



To the 
loads 



Utilization 
voltage 



Transmission Distribution Utilization 

Figure 1.1. —Simple mine electrical system arrangement. 



employing dc concepts. However, ac systems are more com- 
plicated than dc systems and call for meticulous planning; 
if wrong decisions are made, the results can be extremely 
costly in terms of safety, production, and economics. A great 
deal of effort is needed to maintain an electrical power sup- 
ply within the requirements of the individual pieces of 
mining equipment, and mixing ac and dc in the same mine 
has added greatly to the problems. 

This brief review of the development of electrical sys- 
tems in underground coal mines has shown that the mines 
went from minor electrical usage with the introduction of 
rail haulage to almost total dependency on electricity in a 
period of 50 yr. In the same period, surface coal mining 
underwent changes that were as substantial if less numer- 
ous. They were centered around the enormous growth in 
equipment size. 

Surface Mine History 

The first mechanization of strip mining occurred in 1877 
with the application of an Otis-type steam shovel in a Pitts- 
burg, KS, mine (5). 1 This early attempt was somewhat un- 
successful, but it served as an important step in the evolu- 
tion of strip-mining machinery. Several successful attempts 
to use steam shovels and draglines were made in the next 
30 yr, and these proved that the surface mining of coal was 
completely practical. In time, the advantages of electricity 
over steam became more apparent, and the first significant 
introduction of electric-powered shovels was made in the 
early 1910's. 

Whereas dc series motors were universally employed 
in underground rail haulage, the first large motors used 
in surface mining were dc shunt wound because of their 
constant-speed characteristics. These motors almost directly 
replaced the single-speed steam engine found on practically 
all shovels prior to that time and allowed an immediate 
reduction in work force requirements. Before long another 
important advance in shovel design occurred: the applica- 
tion of separate steam engines to power the shovel motions 
of hoist, crowd, and swing. This change gave increased 
flexibility through the individual control of each operation. 
In a short time, the two major shovel manufacturers of that 
era, Marion and Bucyrus, began to produce both steam and 
electric multimotor shovels with similar characteristics (5). 
Since series-wound dc motors had speed-torque relationships 
similar to those of steam engines when they were used for 
this type of duty, they were employed to drive each shovel 
motion. 

The initial distribution for electric shovels was dc be- 
cause of the nature of the power generation, but technolog- 
ical advances soon made ac power systems superior, and 
ac motors were tried with some success. However, by 1927, 
ac-dc motor-generator (m-g) sets and the invention of the 
Ward-Leonard control concept caused these efforts to be 
abandoned. The new control system enabled the motor 
characteristics to be modified as desired within the motor 
and generator commutator limits, and as a result, 
separately excited dc motors became more attractive than 
series-wound motors. The m-g sets functioned as on-board 
power-conversion units, thereby establishing the use of ac 
distribution in surface mines. 

Motor-generator sets driven by synchronous or induc- 
tion ac motors, Ward-Leonard control, and separately ex- 



'Italicized numbers in parentheses refer to items in the list of references 
at the end of this chapter. 



cited dc motors established the standard, and even now the 
combination is used on most mining excavators, especially 
the larger varieties (2). On smaller machines, some single 
ac electric-motor drives with either mechanical-friction or 
eddy-current clutch systems have evolved, but these are 
often driven by diesel engines. 

Present excavating equipment is generally classified 
into three size groups, although actual capacity ranges are 
normally not assigned. Small shovels are used primarily 
in general excavation, while the intermediate types work 
at bench mining and coal production, and large shovels han- 
dle overburden stripping. Draglines of all sizes are used only 
for stripping. Small and intermediate equipment originally 
ran on rails, but crawler mountings that give improved 
mobility made their first appearance in 1925 (5). Today, 
small and intermediate-sized shovels and draglines are 
mounted on two crawlers, while large shovels have eight 
crawlers (2). Large draglines and some intermediate sizes 
are usually walking types that feature a circular base or 
tub that provides low ground-bearing pressure and a walk- 
ing device for mobility. 

The design of surface mine drilling equipment paral- 
leled excavator development. Initially, most drills were 
pneumatic-percussion types, but because electricity was 
readily available in mines, some machines were designed 
with internal motor-driven compressors. By the early 
1950's, large rotary drilling equipment was necessary to 
satisfy the blasting requirements of thick, hard overbur- 
den (5). This drilling equipment was again electrically 
powered and was very successful. 

The most outstanding change that has taken place in 
electrically powered surface mining equipment has been in 
connected horsepower. For example, a 25-yd 3 dragline or 
stripping shovel that had a maximum total load of around 
2,000 hp was considered enormous in the late 1940's (5, 9). 
By 1955, 50- to 70-yd 3 excavators were being manufactured 
with maximum horsepower at 4,650 hp. Five years later, 
shovels had reached a 140-yd 3 capacity wth 12,000 hp of 
main drive motors (7). In 1976, the largest excavator in 
service had 20,000 hp in m-g set drive motors (4). 

Distribution and utilization voltages also increased to 
keep pace with the peak load demands of this machinery. 
Sometimes the mine distribution and machine voltages for 
these excavators remained the same. Until the mid-1950's, 
4,160 and 2,300 V were the usual mine levels (9). Then, with 
the advent of larger concentrated loads, 7,200 V was con- 
sidered advisable (10). However, this level was found to be 
unsatisfactory for the newly introduced machines with a 
capacity larger than 100 yd 3 , and so 13,800-V mine and 
excavator voltage became a standard. With machines hav- 
ing greater then 200-yd 3 capacity, 23,000-V utilization was 
established (4), but even with these substantial increases 
in distribution, some loads up to 1,000 hp continued to be 
driven at 480 V (10). Production shovels with loads up to 
18 yd 3 commonly stayed at 4,160 and 7,200 V, while in 
general, 4,160 V became standardized for machinery with 
1,500 hp or less. As a result, more than one voltage level 
could be required at a mine when excavators of different 
sizes were employed. 

MINE POWER EQUIPMENT 

A few pieces of power equipment have already been 
mentioned but only to the extent necessary to describe the 
concepts of distribution and utilization. The evolution of 



mine systems has resulted in major items of power appa- 
ratus, each serving a specific function (1,9). In general, they 
can be listed as 

• Generation, 

• Main substations, 

• Portable and unit substations, 

• Switchhouses, 

• Distribution transformers and power (or load) centers, 
and 

• Distribution (conductors and connectors). 

The following paragraphs explain these components 
only in sufficient detail for their inclusion in system ar- 
rangements to be understood. More detailed descriptions 
of substations, switchhouses, and power centers are pre- 
sented in chapters 12 and 13, while chapter 8 is devoted 
to distribution. Power generation is beyond the scope of this 
text, but Ehrhorn and Young (13) provide a thorough discus- 
sion of generation related to mining. 

Substations 

It is common mining practice to purchase all or most 
power from utility companies if it is available. As utility 
voltages usually range from 24 to 138 kV, a main (primary) 
substation is required to transform the incoming levels 
down to a primary distribution voltage for the mine. In ad- 
dition to having the transformer, substations contain a 
complex of switches, protection apparatus, and grounding 
devices, all having a function in safety. Main substations 
are often permanent installations. The nature of the min- 
ing operation and its power needs dictate how many main 
substations are required and where they should be placed. 
They may be owned by the utility or the mining company; 
the decision of ownership is commonly dependent on eco- 
nomics. However, if the total connected load is greater than 
1,000 hp, mine ownership is often more favorable (13). 

Portable and unit substations are similar in operation 
to main substations except they serve to transform the 
primary distribution voltage to a lower distribution level. 
The term "unit" means that the substation and power 
equipment are designed and built as a package. In a typi- 
cal strip-mining deployment, a large dragline may require 
24 kV while the production shovels and other mining equip- 
ment need 4,160 V. 

Switchhouses 

Switchhouses are portable equipment that protect the 
distribution circuits. Their internal components are chiefly 
protection devices, with circuit deenergization performed 
by disconnect switches, oil circuit breakers, or vacuum cir- 
cuit breakers. The switchhouse may contain more than one 
complete set of devices, for instance, a double switchhouse, 
which can independently protect two outgoing circuits. This 
category encompasses disconnect switches, which are power 
equipment containing only manual devices, with the prime 
function of allowing mine power to be removed from the 
main supply. 

Power Centers 

At the outermost distribution points there are power 
centers and distribution transformers, which transform and 
convert the distribution voltage to utilization levels. In- 



cluded here are ac to dc conversion equipment or rectifi- 
ers, which convert the distribution voltage to dc for use on 
rail trolley and other systems. The power center, also 
termed a load center, usually implies an internal bus, which 
is defined shortly, in the section, "Radial System." In 
essence, these are all portable substations, and as with 
switchhouses, each outgoing circuit has its own set of in- 
ternal protection components. However, an individual unit 
may supply from 1 to as many as 20 machines. Power 
centers can be considered the heart of an underground min- 
ing section power system. In surface mines, power centers 
supply power to low-voltage auxiliary machinery and loads; 
there may be no need for this equipment with the primary 
mining machinery. 

Distribution Equipment 

This category of major power equipment is often referred 
to as the weakest link in the mine power systems. It en- 
compasses all the overhead powerlines, cables, cable cou- 
plers, and trolley lines used to carry power and grounding 
between the power equipment and eventually to the loads. 
The conductors are usually called feeders when they are part 
of distribution; at utilization, when connected to portable 
mining machines, they are called trailing cables. 

BASIC DISTRIBUTION ARRANGEMENTS 

The basic distribution arrangements available for in- 
dustrial applications are radial, primary selective, primary 
loop, secondary selective, and secondary-spot network (6). 
Radial systems are the most popular arrangements in min- 
ing, though other configurations can be found where special 
circumstances call for greater system reliability (3). Sur- 
face mines have, of course, greater flexibility than 
underground mines and employ a wider range of distribu- 
tion arrangements. Secondary-spot networks, which are the 
most popular system for large facilities in other industries, 
are uncommon but could be applied to preparation and mill- 
ing plants. The following descriptions of the main distribu- 
tion patterns are based on the Institute of Electrical and 
Electronics Engineers (IEEE) definitions (6). This institute, 
the leading national professional electrical organization, 
sets standards and recommended practices that are inter- 
nationally renowned for their correctness. 

Radial System 

Figure 1.2 shows radial distribution in its simplest form. 
Here, a single power source and substation supply all equip- 
ment. The single vertical line represents one connection 
point for all feeders, or all connecting lines, and is termed 
a bus. Voltage along the bus is considered to be constant. 
Radial systems are the least expensive to install as there 
is no duplication of equipment, and they can be expanded 
easily by extending the primary feeders. A prime disad- 
vantage is tied to their simplicity; should a primary com- 
ponent fail or need service, the entire system is down. 

An expanded radial system, the load-center radial, is 
illustrated in figure 1.3. As in figure 1.1, two or more volt- 
age levels are established, but the feeders form a treelike 
structure spreading out from the source. This system has 
the advantages of the simple system and several others too. 
If the load centers or distribution transformers are placed 
as close as practical to the actual loads, most distribution 



Bus 



Substation 



O 

Utility 
source 



Switchhouse 



Switchhouses 



Utilization 
equipment 



Cable or 
powerline 



O 



O 



o 



Source 



Figure 1.2.— Simple radial distribution system. 



Substation 




Circuit 
breakers 



Utilization 
equipment 



Figure 1.3.— Power-center type of radial distribution. 



will be at the higher voltage. This allows decreased con- 
ductor investment, lower electrical losses, and better voltage 
regulation. 

Primary-Selective System 

The primary-selective system (fig. 1.4) adds downtime 
protection through continuity of service. Each substation 
can receive power by switching from either of two separate 
primary feeders. Each feeder should have the ability to 
carry about 80% of the load, so that one feeder can accept 
a temporary overload (10) and provide continued operation 
if one source should fail. During normal service, each 
feeder should handle one-half of the load. The system is 
simple and reliable but costs are somewhat higher than for 
the radial system because of the duplication of primary 
equipment. 

Primary-Loop System 

Though found in some mines, the primary-loop system 
(fig. 1 .5) is not considered good practice. It offers the advan- 
tages and disadvantages of the primary-selective system 
and the cost can be slightly less, but this configuration can 



Main 
Source substation 



Main 
Source su bs totio n 

O 




Unit substations 
Bus — *- 
Circuit breakers — *- 

To the loads 
Figure 1. 4.— Primary-selective distribution system. 



a 



Switchhouse 



Switchhouse 



Sources Substations 

O- 




Loop feeder 
Figure 1. 5.— Primary-loop distribution. 



result in dangerous conditions when a primary feeder fails. 
For instance, a failed portion can be energized at either side, 
creating an extreme hazard to maintenance personnel. 

Secondary-Selective System 

In a secondary-selective system, a pair of substation 
secondaries are connected through a normally open tie cir- 
cuit breaker as shown in figure 1.6. The arrangement al- 
lows greater reliability and flexibility than do the preced- 
ing techniques. Normally, the distribution is radial from 
either substation. If a primary feeder or substation fails, 
the bad circuit can be removed from service and the tie 
breaker closed either manually or automatically. Mainten- 
ance and repair of either primary circuit is possible with- 
out creating a power outage, by shedding nonessential loads 
for the period of reduced-capacity operation. Other methods 
that can be used to provide continuity of service include 
oversizing both substations so that one can carry the total 



load, providing forced-air cooling to the substation in serv- 
ice for the emergency period, or using the temporary over- 
load capacity of the substation and accepting the loss of com- 
ponent life (6). Economics often justify this double-ended 
arrangement if substation requirements are above 5,000 
kVA. Note that the substation capacity or ability to trans- 
form power is rated in kilovoltamperes. 

Secondary-Spot Network 

In the secondary-spot network, two or more distribution 
transformers are supplied by separate primary-distribution 
feeders as illustrated in figure 1.7. The secondaries are tied 
together through special circuit breakers, called network 
protectors, to a secondary bus. Radial secondary feeders are 
tapped to the bus and feed the loads. This arrangement 
creates the most reliable distribution system available for 
industrial plants. If a failure occurs in one distribution 
transformer or primary circuit, perhaps by acting as a load 
to the bus, its network protector can quickly sense the 
reverse power flow and immediately open the circuit. Total 
power interruption can occur only with simultaneous mis- 
haps in all primary circuits or a secondary-bus failure. 
However, this type of system is expensive, and the relia- 
bility gain is not warranted for the majority of mining 
applications. 

It may be obvious that these basic distribution tech- 
niques can be combined into hybrid systems. When this is 
done, there can be confusion about what is primary or 
secondary. Ordinarily, the subsystems are defined by the 
specification of the substations. This will be demonstrated 
in the next two sections. 



UTILITY COMPANY POWER 

As utility companies are the principal power source for 
mines, an understanding of utility system power transmis- 
sion and distribution is important. Often this system greatly 
affects the power available to the mine, including voltage 
regulation, system capacity during power failures in the 
mine, and overvoltage occurrences. 



o 



Sources 



Main 
substations 



Switchhouses 



Normally open 
tie breaker 



o 



Switchhouses 



T TT 



In a nearby substation, power from a generating sta- 
tion is transformed up to a transmission voltage, commonly 
69,000 V or more (6). This power is carried on transmission 
lines to major load areas, either supplying large industrial 
users directly or powering the utility's own distribution 
substations. Distribution substations step the voltage down, 
this time to a primary-distribution level ranging from 4,160 
to 34,500 V, but most often at 12,470 or 13,200 V (6). These 
stages are illustrated in figure 1.8. 

The utility service, therefore, can be any of the follow- 
ing standard values, in kilovolts: 138, 115, 69, 46, 34.5, 23, 
13.8, 12.47, 6.9, 4,800, 4.16, and 2.4 (13). Generally, the 
delivered voltage ranges from 23 to 138 kV, but other values 
such as 480, 2,300, and 7,200 V are also found. What is 
available to the mine depends on whether the possible con- 
nection is to the power company transmission system, a 
primary-distribution system, or a distribution transformer. 

It is the responsibility of the mining company to select 
that voltage best suited to its needs. Primarily, the choice 
depends on the amount of power purchased. It is not safe 
to assume that the power company has the capability 
to serve a large mine complex from existing primary- 
distribution lines or even from the transmission system. The 
problem stems from the fluctuating nature of mine loads. 
For example, large excavators in surface mines can require 
high peak power for a short time, followed by regenerative 
peak power, cycling within the span of 45 s. The fluctuat- 
ing load may create voltage and frequency variations be- 
yond the limit set for other utility customers. Accordingly, 



Sources 



o o o o 



Substations 



Network 
protectors 

Bus 



Switchhouse 



To the loads 
Figure 1. 7.— Secondary-spot network technique. 



Generating 
station 



C^ 



Transmission 
line 



Regulated primary 
/ distribution system 



Substation 



Substation 



Secondary distribution 



Distribution 
transformer 



Figure 1. 6.— Secondary-selective system. 



Branch 
circuit 



Utilization 
equipment 

Figure 1.8.— Representative utility transmission and 
distribution. 



most large draglines and shovels require power from 69- 
to 138-kV transmission systems to get adequate operational 
capacity, and the construction of several miles of trans- 
mission equipment can result in a sizable cost for the mine 
budget. 

Regardless of where the main mine substation is tied 
into the utility complex and who owns that equipment, its 
outgoing circuits will here be termed the mine primary- 
distribution system or just distribution. The incoming power 
will most often be referred to as a transmission system. 

The following sections identify the main types of mines 
in the United States, classify the major equipment em- 
ployed, and describe the power-distribution arrangements 
that are found in them. Of necessity this can be only a very 
brief overview, but it is designed to indicate the problems 
and complexities that can arise in mine power distribution 
and utilization. Individual topics mentioned here are ex- 
panded in detail in later chapters. 

SURFACE MINING 

Surface mining methods are selected over underground 
methods when the overburden, the earth above the coal 
seam, can be removed economically to expose the coal. 
Productivity, safety, and economics usually favor surface 
mining of seams less than 150 ft deep. Surface coal mining 
consists of four basic operations: overburden removal to ex- 
pose the coal, coal loading, haulage, and reclamation. The 
mining method is generally classified according to such 
physical characteristics as topography or land contour, over- 
burden thickness, coal thickness, number of coal seams, type 
of overburden, fragmentation characteristics of the over- 
burden, climate, and hydrology. The mining method is also 
affected by Federal and State requirements. The mining 
method selected must protect the health and safety of the 
workers and minimize environmental disturbance and be 
designed for the specific set of prevailing physical condi- 
tions. The major surface mining methods for coal are con- 
tour mining and area mining. 

Contour mining methods are commonly used in rolling 
or mountainous terrain; they are called contour mining 
because overburden removal progresses around the hillside 
at the coal seam horizon such that the pit resembles a con- 
tour line. There are many varieties of contour mining, but 
in all methods overburden is fragmented by drilling and 
blasting, and removed to expose the coal seam. The over- 
burden may be removed by small diesel or electric drag- 
lines, or by diesel-powered front-end loaders and trucks. In 
soft overburden and for topsoil removal, scrapers and bull- 
dozers may be used. 

Area mining is the predominant stripping method in 
more level terrain. As its name implies, area mining can 
cover an extensive region, using various box-cut or strip pit 
and benching techniques. It may be used to mine both thick 
and thin seams, or multiple seams; where these seams are 
dipping, area mining is modified to approximate the open 
pit methods common in metal mining. In all cases, over- 
burden handling and reclamation are an integral part of 
the process. Equipment varies, depending on the scale 
of the operation, from small draglines and dozers to 
massive equipment that has more than 30,000 connected 
horsepower. 

In general, the magnitude of electrical distribution and 
utilization is greater in area mining than in contour min- 
ing. Combination of equipment employed in large multi- 
seam operations may include tandem draglines, dragline 



and shovel, pan scrapers with attendant dozers, and drag- 
line and bucket-wheel excavators. Bucket-wheel excavators 
can be very effective where overburden is soft and does not 
require drilling and blasting. Front-end loaders, electric and 
diesel shovels, ripping dozers, and tracked highlifts can all 
be combined with truck haulage for coal removal. 

POWER SYSTEMS IN SURFACE MINES 

Mine power systems can be divided into three categories, 
depending upon the purpose of a specific portion: 

1. Subtransmission, 

2. Primary distribution (or distribution), and 

3. Secondary distribution (or utilization). 

Often, if a subtransmission system is needed, it will have 
the same general arrangement in any mine. At distribu- 
tion and utilization, power-system installations can vary 
greatly, but in some mines distribution and utilization can 
be the same system. Electrical installations in surface coal 
mines are regulated under 30 CFR 77 (14). 

Main Substations and Subtransmission 

Main substations may range from 500-kVA capacity, 
supplying 480 V for only pumps and conveyors, to 50,000 
kVA, servicing a large strip-mining operation and prepara- 
tion plant (10). The substation location is usually an eco- 
nomic compromise between the cost of running transmis- 
sion lines and power losses in primary distribution. From 
the main substation, power is distributed to the various 
centers of load in the operation. However, individual loads 
or complexes, such as preparation plants and other surface 
facilities, may have large power requirements or be so iso- 
lated that primary-distribution operation is not practical. 
In these cases, or for safety reasons, incoming utility trans- 
mission should be extended close to the load. The extension 
is designated a subtransmission system, and the conductors 
are usually suspended as overhead lines (13). 

As shown in figure 1.9, subtransmission commonly re- 
quires a primary switchyard of high-voltage switching ap- 
paratus for power tapping. Branch circuits are fed through 

Incoming transmission lines 
Y 



Primary 
switchyard 



t 



Normally open L j J Dual 
tie breaker / bus 

i — i 




Subtransmission 



Preparation 
plant substation 



To preparation 
plant 



1^ Second 
-< — \y subtransmission 
if primary 
selective or 
secondary 
selective desired 
on major load 
concentration 



To mine distribution 
Figure 1. 9.— Subtransmission for surface mine. 



circuit breakers to protect the subtransmission line and the 
utility's system. Dual-bus configurations are employed if 
primary-selective or secondary-selective distribution is de- 
sired on major load concentrations to provide high reliabil- 
ity. This additional subtransmission circuitry is illustrated 
in figure 1.9 by dashed lines. 

Subtransmission circuits, primary switchyards, and 
main substations are almost always permanent installa- 
tions located in areas unaffected by the mining operation. 
The main substation is where the grounding system for the 
mine is established. This ground is carried along the 
powerlines through overhead conductors or in cables and 
is connected to the frames of all mobile mining equipment. 

Surface Mine Distribution 

Mine power distribution, in its simplest radial form, has 
already been shown to consist of a substation, distribution, 
and a power center feeding the mining equipment. The ar- 
rangement is very common in small surface operations 
where the distribution voltage is commonly 4,160 V but 
can be 2,300 V in older equipment. In the smallest mines, 
power is purchased at low-voltage utilization (often 480 V) 
and fed to a distribution box to which motors and equip- 
ment are connected. At times, simple radial distribution is 
employed in large surface mines where only one machine 
must be served or an extensive primary-distribution net- 
work cannot be established, as in some contour operations. 

The great majority of strip mines employ radial distribu- 
tion, but secondary-selective and primary-loop designs can 
also be found. Simplified examples of the three systems are 
provided in figure 1.10 to 1.12. In all configurations, a por- 
tion of the primary distribution is established at a base line 
or bus. The base line is usually located on the highwall, 
paralleling the pit for the entire length of the cut. Its loca- 
tion is typically maintained 1,500 ft ahead of the pit, and 
it is moved as the pit advances (3). Distribution continues 
from the base line to the mining equipment, with the con- 
nections maintained at regular intervals. As the machines 
move along the pit, the base-line connections are changed 
to another convenient location. 

The base line can consist of overhead polelines or a 
cable-switchhouse configuration, figures 1.13 and 1.14 (3). 
It can be seen that cable distribution brings power into the 
pit area, where shielded trailing cables connect to the ma- 
chines. The overhead poleline plus cable arrangement is 
common in older mining operations, especially when utiliza- 
tion is at 7,200 V or less (3). Typical spacing between poles, 
or line span, is 200 ft. Drop points are noted in figure 1.13 
by triangles. These are terminations between the overhead 
conductors and the cables, mounted about 8 ft above the 
ground on poles spaced at regular intervals of around 1,000 
to 1,500 ft. 

Cables connected to the drop points deliver power to 
skid-mounted switchhouses located on the highwall or in 
the pit. The switchhouses may contain manual disconnect 
switches, which are commonly termed switch skids or 
disconnect skids, automatic circuit-protection devices or 
breaker skids, or a combination of both. The skids can either 
be boat design with flat bottoms or have fabricated runners, 
depending upon the allowable bearing pressure of the mine 
terrain. Couplers or plug-receptacle pairs are commonly 
used for both feeder and trailing-cable connections. Discon- 
nect and circuit-protection functions are required for each 
distribution load, and double switchhouses (two-breaker 
skids) are frequently employed for two loads. Unit substa- 



tions often contain internal circuit protection on the incom- 
ing side, and thus do not require a breaker skid. 

Trailing cables are usually 1,000 ft in length, although 
lengths to 2,000 ft can be found. When longer cables are 
necessary to reach a breaker skid, in-line coupling systems 
can be used, and these are commonly mounted on small 
skids for easy movement. Trailing-cable handling for strip- 
ping equipment is often assisted by cable reels mounted on 
skids or self-propelled carriers. Large excavators can require 
the self-propelled variety. 



Switchhouses 



1_ 



Unit 
substation 



Base line ■*■ 



Pit highwall 



9~Q 



h^H 



-T] [I ►Base line 

Dragline 




Production 
shovel 

Pump, lighting 
Figure 1.10.— Radial strip mine distribution system. 



Main 
substation 



I 



Base line 



— a— Lp 



Normally open 
tie breaker 

/ 



I 



Main 
substation 



. Switchhouses 




Pit 
highwall 



Production 
shovels 



Other pit 
power 



Figure 1.11.— Secondary-selective distribution in strip min- 



ing. 



Main 
substation 




Figure 1.12.— Primary-loop design for strip mining. 



10 



Utility company 
metering 



7. 2- kV overhead poleline 
( base line ) 

1,000 ft /1,000 ft 1,000 ft 1,000 ft 
•\ 1 — \ — i *\/ rAr 




Power company supply, 
69 kV 



Substation 

69kV/7.2kV) 



V V V 

Lateral 
cables 



Trailing 
cable 



Dragline 



To 

auxiliary 
equipment 

Production 
shovel 
Drill 



7.2-kV 
to 4,160-V 
trans- 
former^ 

TBS 

Htbs r— ' 

Trailing 
cable 

Dragline 

Production 

shovel 

Spare 

KEY 

DS Disconnect switch 

TBS 2- breaker skid 




Lateral 
'Cable 
7.2-kV 
to 
\4,160-V 
■ trans- 
1 former 



v 



TBS 



Drill 



Figure 1.13.— Radial distribution for strip mine with 
overhead-poleline base line. 



Utility company supply, 
138 kV 



Substation 
(!38kV/7.2kV) 



Additional J 
distribution -^ 
if pullback 
mining method 
is used 

7.2-kV cable 
by incline to 
pullback machine 



138-kV 
powerline 



] Utility company 
metering 

Overload circuit breaker, 
138 kV 



Substation 
(138kV/25kV) 

25-kV cable 



Skid-mounted 
triple - switch 
skids 



Single-breaker skid 
Dragline 25 . kV \ 

(transformed trailing cable i — i 
to 7,200V - — — I \~ 1 1 

on machine ) Lateral cables/ 



2-breaker skid 



r? 



t 

Dragline ! 



Single-breaker skid 

Spare ■* 

25 kV/ 440 V 
Auxiliary 440-V coble 



tl 






1,500 ft 



25-kV 



^Of'Kcoble 



Production 
shovel 



equipment 
2-breaker skid 

Drill- 

Spare- 1 



o- 



Lateral cables-^ 



1 



7.2-kV cables 
25-kV/ 7.2-kV transformed 

Drill -. 



•1,500 ft 



1,500 ft 



Trailing cables 
Production shovel — — 



2-breaker 
skid 



Figure 1.14.— Radial distribution for strip mine with all-cable 
distribution. 



The layout for an all-cable mine distribution, figure 1.14, 
is very similar to that just described. In this case, how- 
ever, the base line is assembled using cable-interconnected 
switchhouses. As noted in the illustration, the common 
approach is to use disconnect skids with three internal 
switches in the base line and to have separate breaker skids 
in line with the cables feeding the machinery. Another ap- 
proach is to combine the single-breaker skids into the base- 
line switchhouses. 

When a secondary-selective configuration is used, as 
shown in figure 1.11, a normally open tie circuit breaker 
is placed in the base line in a location approximately equi- 
distant from the main substations. In some operations, the 
two substations and the tie circuit breaker may be in the 
same location, with two feeders running from the substa- 
tion area to the base line. More than two main substations 
may be established in very large operations. 

Primary-loop systems have occasionally been used in 
strip mining. It can be noted from figure 1.12 that the sub- 
stations actually operate in parallel, considering the base 
line to be a bus. Here the substations can be smaller than 
those needed for a radial system. Notwithstanding, certain 
precautions should be taken with this configuration (9). 
For example, the substations must be identical if they are 
to share the load, but as an unbalanced load distribution 
is probable on any system, it is likely that the two substa- 
tions will not be equally loaded. Regardless, because of the 
safety hazards, primary -loop distribution is considered un- 
satisfactory and is not recommended. 

Distribution voltage for the surface mine may be 7.2, 
13, or 23 kV, and to a lesser extent 4.16 kV. Regardless 
of the level, drills and production shovels usually operate 
at 7,200 or 4,160 V. Therefore, when higher distribution 
levels are needed, portable unit substations are commonly 
used in the pit. One instance would be when the load cre- 
ated by a large machine is several times that for auxiliary 
machines. Another method is to establish two base lines 
on the highwall for two distribution voltages, as shown in 
figure 1.15. Here, a large unit substation interconnects the 
two base lines. Even in this situation, as can be seen in 
the preceding illustrations, low-voltage unit substations 



Utility 



Main 
substation 



CbFD — a — — D 




Switchhouse 
for 23-kV 
base line 



Switchhouse 
for 7,200-V 
base line 



Dragline 



10,000-ft pit 



Figure 1.15.— Surface mine distribution system using two 
base lines. 



11 



or power centers are often required for 480-V auxiliary 
equipment. 

The primary purpose of any primary-distribution 
scheme in a surface mine is to provide a flexible, easily 
moved or modified power source for the highly mobile 
mining equipment. System designs must also be considered 
as an integral part of the total mine operation. Those 
described have these objectives in mind. As will be seen, 
the distribution system in any surface or underground mine 
that serves portable equipment is subject to damage from 
the mining machinery itself, and as a result, the system 
must be designed with optimum flexibility and considera- 
tion for personnel safety. 

Open pit power systems are quite similar to those in 
stripping mines but with one main exception: primary dis- 
tribution typically establishes a ring bus or main that 
partially or completely encloses the pit. Radial ties to the 
bus complete the circuit to switchhouses located in the pit, 
and portable equipment again uses shielded trailing cables. 
An example is shown in figure 1.16. Distribution voltage 
is normally 4.16 kV, but 7.2 or 6.9 and 13.8 kV are some- 
times used. Unit substations are employed if equipment 
voltages are lower. Primary distribution is almost invari- 
ably through overhead lines. 

UNDERGROUND COAL MINING 

Figure 1.17 is a plan view of a typical U.S. underground 
coal mine. A system of main entries, each 16 to 20 ft in 
width, is developed from the coal seam access point to the 
production areas, which are called panels or sections. Pil- 
lars of coal are left during mining to support the overburden 
above the entries. Crosscuts are mined between the entries. 
The main entries may remain standing for several years 
while coal is being extracted from the panels. Haulage of 



Utility 



Substation 



Disconnects 
\ 



Ultimate pit 
limit 




Substation 



Utility or 
subtransmission 



Overhead line 
ring main 



Poles 



Substation 



Utility or 
subtransmission 



personnel, supplies, and coal, together with provision of 
ventilating air and dust-suppression water, and electrical 
distribution are necessary functions of the main entries 
throughout the life of the mine. 

The mining method is defined by the configuration of 
the open workings and by the classification of equipment 
used. The important underground coal mining methods are 
room and pillar, which may be either conventional or con- 
tinuous, and longwall. To the miner, the type of mining 
machinery used is implied by each category. The room-and- 
pillar method remains dominant in the United States, 
although there has been a recent substantial increase in 
longwall mining. The choice of a specific mining method 
is frequently dictated by such natural conditions of the mine 
as the characteristics of the overburden, roof, and floor, plus 
the seam dip, water, methane, and seam height (11). Es- 
sentially, the method and equipment selected are based on 
the combination that will provide the safest and most prof- 
itable extraction within the given set of geologic conditions, 
while complying with State and Federal health, safety, and 
environmental regulations. 

Room-and-Pillar Mining 

Room-and-pillar mining is named for the regular pat- 
tern of openings made in the coal seam and was the earliest 
form of underground coal workings. 

Conventional Mining 

The conventional mining method represents a direct 
evolutionary link with the early mining techniques. It is 
based on the original loading machine, which came into use 



Pillars 



84 ft 



Previously mined 

longwall panel, 

roof caved (Gob) 




Room-and-pillar 
panels developed 



Panel 
entries' 



Crosscuts ' 



Longwall 
face 



Submain 
entries 



Main entries 



Figure 1.16.— Open pit power system. 



Figure 1.17.— Layout of underground coal mine. 



12 



in the early 1920's. Modern conventional mining consists 
of six distinct operations: 

1. Undercutting the coal face, 

2. Drilling holes in the face for blasting, 

3. Blasting, 

4. Loading the broken coal onto a face haulage system, 

5. Hauling the coal from the face area to a subsequent 
haulage system, and 

6. Providing roof support. 

In order, these steps comprise a mining cycle; after roof sup- 
port is completed, work begins again at step 1. Ventilation, 
although essential, is not included as a separate step in the 
cycle as it must be provided continuously. Other safety pro- 
cedures include careful examination of the face and roof 
after blasting and before each job begins at the face. 

Mobile self-propelled mining equipment performs most 
of the operations in conventional mining. The cutting ma- 
chine, basically an oversized chain saw, is employed to cut 
a slot at floor level, called the kerf, which allows coal ex- 
pansion during blasting. A self-propelled face drill follows 
the cutter and makes several holes in the face with its 
carbide-tipped auger-type drill bits. Blasting is carried out 
either by chemical explosives approved as permissible by 
the U.S. Mine Safety and Health Administration (MSHA) 
or to a lesser extent by high-pressure air. Permissible ex- 
plosives will not ignite methane and coal dust when used 
correctly. 

A crawler-mounted loading machine loads the broken 
coal onto the face haulage vehicle, typically a shuttle car. 
The car is equippped with a chain conveyor that moves the 
coal from the load end to the discharge end and subse- 
quently unloads it from the vehicle. Shuttle cars almost 
invariably work in pairs and move the coal to rail cars or 
a conveyor belt, which makes up the next stage of materi- 
als handling in the mine. The roof bolter, sometimes called 
a roof drill, is a rubber-tired vehicle that secures the roof 
by first drilling vertical holes and then emplacing roof bolts, 
which secure the roof either by clamping thin roof layers 
together to form a thick beam, or by hanging weak strata 
to a more competent upper layer. Drilling is usually ac- 
complished by rotary action with auger-type bits. The re- 
sulting dust is collected through the bit and hollow drill 
rod by vacuum. 

With few exceptions, all these machines are electrically 
driven, powered via trailing cables from the mine power 
system. Since the mining equipment is continually moved 
among several faces in a coordinated plan designed for 
maximum production efficiency, the handling of trailing 
cables is a significant part of the mining cycle. The result 
of coal removal is a system of open rooms divided by coal 
pillars that support the roof as mining advances toward the 
property boundaries. When the equipment approaches the 
property limit, the operation is turned around and retreat 
mining takes place. If surface subsidence is permitted, the 
pillars are removed in an organized extraction plan and the 
roof is allowed to cave. The broken material that then fills 
the mined void is known as gob. 

Continuous Mining 

The heart of the continuous coal mining method is the 
continuous mining machine, which replaces the conven- 
tional room-and-pillar unit operations of cutting, drilling, 
blasting, and loading. The mining functions of haulage and 
roof support remain, although some continuous miners also 



perform roof bolting. The term "continuous" is actually a 
misnomer because of legal constraints that mandate inter- 
ruptions in the mining process for safety checks and ven- 
tilation requirements. 

The most common form of face haulage in continuous 
mining is again the shuttle car. One of the main problems 
associated with continuous mining is the intermittent 
nature of the shuttle car haulage system, which causes fre- 
quent delays at the face. As a result, various types of con- 
tinuous haulage systems have been developed to alleviate 
this problem. Mobile chain and belt conveyors are the most 
popular of these systems, and these are applicable to min- 
ing low coal. Continuous haulage systems have not been 
without problems, and some designs have been hampered 
by poor reliability and lack of maneuverability. Hydraulic 
systems have shown great promise; they operate by pulver- 
izing and slurrying the coal immediately behind the miner, 
then pumping the slurry to the surface. 

Longwall Mining 

Longwall mining is the most popular underground coal 
mining technique in Europe, and it is growing rapidly in 
the United States. In contrast to room-and-pillar mining, 
longwall is capital intensive rather than labor intensive. 
Longwalls are usually 300 to 600 ft wide, and the direction 
of mining with respect to the main entries classifies them 
as either advancing or retreating longwalls. The latter is 
the most frequent in the United States. 

A typical retreating longwall is shown in figure 1.18. 
The section of coal to be mined, the longwall panel, is first 
delineated by two room-and-pillar entries or headings 
driven perpendicular to the main entry. These two head- 
ings, the headgate and the tailgate, handle haulage equip- 
ment and ventilation. The longwall panel is then mined 
back and forth, retreating toward the main entry. The roof 
is allowed to cave immediately as the longwall equipment 
moves, as is shown by the gob area on the diagram. 

The longwall equipment consists of an interconnected 
system of cutting machine, roof support equipment, and 







Direction 
of mining 



V 



Headgate 
entries 



Longwall 
panel 



Barrier pillar 



Tailgate 
entries 



Main 
entries 



Figure 1.18.— Plan view of retreating longwall. 



13 



conveyor haulage. The cutting machine moves along the 
face on a conveyor that also carries away the mined coal. 
Behind the face conveyor, and connected to it, is the roof 
support equipment, which supports a protective metal can- 
opy or shield that extends over the face area. These roof 
support units provide both the protection and the forward 
mobility of the system. 

The typical face conveyor is a flexible armored-chain 
conveyor powered by motors at the headgate and tailgate. 
Mined material moves toward the headgate, where it dis- 
charges to the panel belt via an elevated intermediate 
haulage unit, the stage loader. 

Shortwall mining is a less common mining method; it 
is very similar to longwall mining except that the short- 
wall panel is normally 150 to 200 ft wide. From the stand- 
point of equipment, shortwall can be considered as a com- 
promise between room and pillar and longwall in that the 
extractive and face haulage systems are identical to those 
in continuous mining, while the roof support equipment is 
similar to that used in longwall mining. 



POWER SYSTEMS IN UNDERGROUND MINES 
Regulations 

Underground mine power systems have different char- 
acteristics from those for surface mines, and these two basic 
mining operations are regulated by separate codes and 
standards. For instance, although 30 CFR 77 covers elec- 
trical installations of surface coal mines and surface facili- 
ties of underground coal mines, Part 75 regulates the under- 
ground installations and Part 18 specifies standards for 
electrically powered face machinery (14). Part 77 illustrates 
an overlap between surface and underground legal de- 
mands, which is logical because the surface electrical 
counterparts of both mine types are similar; examples in- 
clude substations and subtransmission. Figure 1.19 can be 
compared with figure 1.9 to see the similarity between sur- 



Incoming transmission lines 



Primary 
switchyard 



6 6 6 6 6 



Note: 

Dual- bus 
configuration 
can be used if 
second source 
desired 



To fan 
power 
system 



To 

surface 
loads 



Mam 
substation 1 



Switches 

and circuit 

breakers 



1 
To 
preparation 
plant 

Main 
substation 2 



Subtransmission 

Preparation 

plant 
substation 



Borehole 1 
to underground 



Borehole 2 
to underground 



Figure 1.19.— Subtransmission for underground mine. 



face mine and underground mine subtransmission. As a 
general situation, however, the mine distribution system 
is related to the mining method. Hence, underground mine 
systems become different from surface mines at the point 
where the circuits leave the substation and go underground. 

Underground Mine Distribution 

As shown in figures 1.20 and 1.21, underground mine 
power systems are somewhat more complicated than those 
for surface applications. Because of the nature of the mine 
and its service requirements, distribution must almost 
always be radial (fig. 1.20); the freedom in routing distri- 
bution enjoyed by surface mines is not available under- 
ground. For increased reliability, secondary-selective main 
substations are employed (fig. 1.21). The secondary-selective 
operation is defined by the use of two substations and mine 
feeders with a normally open tie breaker. Primary-distri- 
bution voltage is most commonly 7,200 V; however, older 
4,160-V systems can still be found, and 12,470 V has in- 
creased in popularity in recent years, especially for long- 
wall operations. The grounding system for the underground 
mine distribution must be separated from that used for sur- 
face equipment. 

Power and mine grounding are fed underground in in- 
sulated cables, either through a shaft or borehole or a 
fresh-air entry. The cables terminate in disconnect switches 
within 500 ft of the point of power entry into the coal seam. 
These switches allow total removal of underground power 
in an emergency. From the disconnects, which may be a 
part of a switchhouse, the power is distributed through 
cables to power centers or rectifiers located as close to the 
machinery as practical. All the cables on high-voltage cir- 
cuits, usually involving only distribution, must have shield- 
ing around each power conductor. 

The prime load concentrations in underground mining 
are created by the mining sections. Distribution terminates 
at the section power center, which is a transformer com- 
bined with a utilization bus and protective circuitry. From 
this, several face machines are powered through couplers 
and trailing cables. Power-system segments for typical con- 
tinuous and longwall operations are given in figures 1.22, 
1.23, and 1.24. Rated machine voltage for most installations 
is 550 Vac, but 250- Vdc and 440-Vac equipment is used 
extensively, and 950 Vac has become quite popular for 
high-horsepower continuous miners and longwall shearing 
machines. In the longwall system, power is fed through 
controls to the various motors. On conventional or con- 
tinuous equipment, the utilization approximates the radial 
system shown in figure 1.22. 

If belt haulage is used, distribution transformers are 
located close to all major conveyor belt drives and are re- 
ferred to as belt transformers. After transformation, power 
is supplied through starter circuitry to the drive motors. 
With rail haulage, distribution terminates at rectifiers that 
contain a transformer and rectifier combination. The rec- 
tifiers are located in an entry or crosscut just off the rail- 
way. As shown in figure 1.25, dc power is then supplied 
through circuit breakers to an overhead conductor or trolley 
wire and the rail, with additional rectifiers located at reg- 
ular intervals from 2,000 to 5,000 ft along the rail system. 
For further protection, the trolley wire is divided into elec- 
trically isolated segments. The typical rectifier supplies the 
ends of two segments of trolley wire and each feeder has 
its own protective circuitry to detect malfunctions. Each 
trolley-wire segment is called a dead block. This loop-feeding 



14 



arrangement is continued throughout most of the haulage 
system except for the most inby segment, which is dead- 
ended. In some mines, dc face equipment and small dc 
motors are powered from the trolley system through a fused 
connection (or nip) to the trolley conductor and rail. The 
dc distribution can also serve large motors directly through 
switchgear; however, this practice is rare in underground 
coal. 



All power equipment used underground must be rugged, 
portable, self-contained, and specifically designed for in- 
stallation and operation in limited spaces. In addition, all 
equipment and the cables connecting them must be pro- 
tected against any failures that could cause electrical haz- 
ard to personnel. This is primarily provided by protective 
relaying built into each system part, with redundancy to 
maximize safety (4). 



Utility company I I 
metering UJ 



Main 
substation 



Ground 
level 



Rectifier 



Cool 
seam 



To other 
distribution 




Miscellaneous loads 
(shop, pump, etc.) 



CONTINUOUS 
MINING SECTION 



IJUUUUUUUU 

/|\ /Tv /K /T\ /K /K /K /t\ 



3 Oj — (M 

° s t « 



•t- •« 
o a 



S\ /\ /\ S\ /\ 



01 o • 



/f\ /*\ y^ /K /K 



Master 
control 



e e - 



O -O TJ 
W >* >s 

z r 



Figure 1.20. -Radially distributed underground power system. 



15 



MAIN 

SUBSTATION 

AREA 



Ground 
level 

Coal 
seam 



r 


















rrr t 




~*— Substations 




'"# 1 


h r 


4r 


" Normally open 
tie breaker 



Shafts or boreholes 






To other portable ^_ 
switchhouses and 
loads (1/2 total) 



-» 






Disconnect switchhouses 



^>-r-»- 

UJ 



UJ 



To other portable 
switchhouses and 
loads (1/2 total) 



Major load concentrations 
Figure 1.21. -Secondary-selective distribution In underground mines. 



Low -voltage couplers 



High-voltage 
couplers 




Feeder 
cable 

/ 
C 




Trailing 
cable 

iter 


Continuous miner 






"5 

r 










Bolter 
Shuttle car 
Shuttle car 
s Feeder 






L 




pow 


Mine 
/er co 



Figure 1.22. -Utilization in continuous mining section. 



16 




/ 

2 

3 
4 
5 

6 
7 
8 



Self-advancing supports 



Stage loader 



KEY 
Motor, 125 hp 
1,000- kVA power center 
125 -hp stage- loader starter 
Dual 125-hp face-conveyor starter 
Dual 75- hp pump and 230 -hp shear starter 
Pump, 75-hp 
Pump, 75-hp 
Master control 

Figure 1.23. — Power-system segment with longwall equipment. 



High-voltage feeder cable 

High-voltage coupler 



Medium-voltage cables 



Permissible 

medium-voltage 

couplers 



Power center 



Medium- voltage outlets 



mm 



Face -conveyor 
starter ^ 



un 



uu^ 



Medium- voltage couplers 

Medium-voltage cables 



Master 
control 



Cables for control 
of starters 



/ 



d_d 



u~u 



II II 



D_a 



UUu" 



Shearer 
and stage- 
loader starter 



d_a 



Medium-voltage 
cables 



if \ 



i 



UUP" 



IMMI 



Permissible 

medium -voltage 

couplers 



Hydraulic- pump 
starter 



Medium-voltage 
cables 



950 -V face 
conveyor motors 



950-V shearing- ,950-V stage- 950-V hydraulic- 
machine motors loader motor pump motors 



Figure 1.24. -Diagram of electrical-system segment for longwall. 



17 



High- voltage feeder 



Switchhouse 



Switchhouse 



Power from substation 



«n-«r 

UJ 



IT 



'/f\~ 



High-voltage 
ac input 



i 1 

«-r-«- 

LaJ 



"4\ 



A 

r 



/" 



Low -voltage 
dc output 



Rectifier 



2 positive outputs 
to trolley conductor, 
each protected by a 
separate rectifier 



A 



■*• To other downstream 
switchhouses 



Rectifier 




Dead-block 
segment 

Figure 1.25. -Parallel-feed haulage system. 



Dead - ended 
segment 



SURFACE FACILITY POWER REQUIREMENTS 

The surface activities of any mine, which may include 
shops, changing rooms, offices, ventilation fans, hoisting 
equipment, preparation plants, and so forth, can have large 
power requirements. For safety, these facilities should at 
least have an isolated power source and at times a separate 
substation. 

In preparation plants, the distribution arrangements 
are almost always expanded radial or secondary selective 
(8). Representative system layouts are shown in figures 1.26 
and 1.27. In both, distribution is at 2.4 to 13.8 kV, with 
4,160 V the most common level. Power is distributed at one 
of these levels to centers of electric load. This power may 
be used directly for high-voltage motors, but usually the 
voltage is stepped down to supply groups of motors or single 
high-horsepower motors. The power centers must be in an 
elevated location or totally enclosed. The rooms used for 
these and other electrical components may also be pressur- 
ized to exclude coal dust. 

The most popular voltage for preparation plant utiliza- 
tion is 480 V. This voltage is used to drive all motors 
throughout the plant except those with high-horsepower 
demands, such as centrifugal dryers and large fan drives, 
where 2,300 or 4,160 V is commonly employed. These higher 



voltages may also be preferred for any motor that requires 
continuous service or independence from the power-center 
loads. Note that 240-V motors are unsatisfactory for typi- 
cal preparation plant demands. Most modern preparation 
plants use group motor control instead of individually 
housed control units, since this method facilitates main- 
tenance and enables the interlocking of the various motor 
functions required for semiautomatic facilities. All manual 
controls, indicating lights, and so on are grouped in one 
central operating panel to allow easy access and visual 
indication of plant operation. The panel is often called a 
motor control center (MCC), as shown in figure 1.27. 



BASIC DESIGN CONSIDERATIONS 

The goal of the power engineer is to provide an effic- 
ient, reliable electrical system at maximum safety and for 
the lowest possible cost. The types of information made 
available to the power engineer include the expected size 
of the mine, the anticipated potential expansion, the types 
of equipment to be used, the haulage methods to be em- 
ployed, and whether or not power is available from a util- 
ity company. The amount of capital assigned for the elec- 
trical system will also be designated. 



18 



From utility to 

subtransmission 

system 



24 to 13.8 kV 



Substation 



[;] □ — a □ a 

i 





I 

— 1 — 




y 


Power 




center 






Y ^ 


I 


T 



Power centers 



Thermal 
dryer 



Loadout 
station 



Slope 
conveyor 



_mrtL pry 

" T T I T T T T I 

To 480-V preparation plant loads 



Figure 1.26.— Representative expanded radial distribution 
for preparation plant. 



The designed system must meet certain minimum 
criteria. IEEE (12) has defined these basic criteria for in- 
dustrial electrical systems that must be applied to mines: 

• Safety to personnel and property, 

• Reliability of operation, 

• Simplicity, 

• Maintainability, 

• Adequate interrupting ability, 

• Current-limiting capacity, 

• Selective-system operation, 

• Voltage regulation, 

• Potential for expansion, and 

• First cost. 

Of these, safety, reliability, and simplicity are closely re- 
lated. All are dependent on good preventive maintenance. 
In the cramped uncompromising environment of an under- 
ground mine, these are of vital concern. Since continuous 
operation is the aim of every mine operator, planned main- 
tenance should be held to a minimum. Most routine main- 
tenance should be capable of being performed by unskilled 
personnel, since it will be done by the miners themselves. 
Training for these tasks must be provided. 

Adequate interrupting capacity, current-limiting capa- 
bility, and selective-system operation are projected at safety 
through reliability. The first two areas ensure protection 
during a disturbance. Current limiting, when applied to 
grounding, is perhaps the most significant personnel safety 
feature of mine electrical systems. Selective-system opera- 
tion is a design concept that minimizes the effect of system 



Utility metering 




Motor 
control center 



Motor 
control 



Motor 
control center 

•i/ \l ~ i/ 



in w i in 



480-V motors 



4,160-V 
crusher 
motors 



480-V 
motors 



Motor 
control center 

^ vl/ V \b 

nil 

About 2,500 -hp, 
4,160-V motors 



Motor 
control center 

\b 4> 4< ™ 

IIII 

About 2,500-hp, 
4,160-V motors 



Motor 
control center 

-J-j-j-J 



Motor 
control center 

_\|/ \|/ \ l \L \l/ -is 

in in 



480-V motors 480-V 

motors 



Row-coal circuit 



Coarse- 
coal 
circuit 



Pumps, conveyors, blowers, 
filters, fans, jigs, etc. 



Fine-coal 
c i re u i t 



Auxiliaries 



Figure 1 .27. — Representative secondary-selective distribution for preparation plant. 



19 



disturbances. Voltage regulation is a limiting factor in 
system design, particularly underground, and is often the 
main constraint to system expansion. It should be antici- 
pated that when the size of the mine is increased, this might 
involve augmenting the power-system supply through ad- 
ditional power sources. 

While first cost is important, it should never be the 
determining factor, since high-cost equipment, projected at 
maximizing safety and reliability, can easily offset the in- 
creased first cost through the reduction in operating costs. 
At times, this fact appears to elude some company pur- 
chasing agents. 

Using the data available, the task of the power engi- 
neer is to select one combination of power equipment over 
another, provide power or circuit diagrams, estimate the 
equipment, operating and maintenace costs, set the speci- 
fications for the system, and receive and assess the proposals 
from suppliers. For success, the engineer requires a firm 
knowledge of mine power systems, but this understanding 
cannot be based on a "standard mine electrical system" 
because such a standard does not exist: no two mines are 
exactly alike. The engineer must resort to the fundamen- 
tal concepts, an awareness of what has worked in the past, 
and a clear understanding of the legal constraints. This in- 
formation is provided in the subsequent chapters. 

REFERENCES 

1. American Standards Association. American Standard Safety 
Rules for Installing and Using Electrical Equipment in and About 
Coal Mines (M2.1). BuMines IC 8227, 1964. 



2. Bergmann, R. W. Excavating Machinery. Ch. in Standard 
Handbook for Electrical Engineers. McGraw-Hill, 10th ed., 1968. 

3. Bucyrus-Erie Co. (South Milwaukee, WI). Surface Mining 
Supervisory Training Program. 1976. 

4. Cranos, J. C, and D. E. Hamilton. Portable Substations for 
Mine Power Systems. Ind. Power Syst, v. 19, Mar. 1976. 

5. Hollingsworth, J. A., Jr. History of Development of Strip 
Mining Machines. Bucyrus-Erie Co., South Milwaukee, WI, 1967. 

6. Institute of Electrical and Electronics Engineers (New York). 
Recommended Practice for Electric Power Distribution for In- 
dustrial Plants. Stand. 141-1986. 

7. Jackson, D., Jr. Coal Mines. Ch. in Standard Handbook for 
Electrical Engineers. McGraw-Hill, 10th ed., 1968. 

8. Lordi, A. C. Electrification of Coal Cleaning Plants. 
Mechanization, v. 20, Oct. 1956. 

9. Trends in Open-Pit Mine Power Distribution. Coal 

Age, v. 66, Jan. 1961. 

10. Rein, E. C. Electrical Apparatus for Surface Mining Opera- 
tions. Ch. in Surface Mining. Soc. Min. Eng. AIME, 1968. 

11. Robinson, N., II. Underground Coal Mining Equipment. Ch. 
in SME Mining Engineering Handbook. Soc. Min. Eng. AIME, v. 1, 
1973. 

12. Stefanko, R. Coal Mining Technology Theory and Practice. 
Soc. Min. Eng. AIME, 1983. 

13. Thuli, A. J. Power. Sec. in SME Mining Engineering Hand- 
book, ed. by J. M. Ehrhorn and D. T. Young. Soc. Min. Eng. AIME, 
v. 2, 1973. 

14. U.S. Code of Federal Regulations. Title 30 -Mineral 
Resources; Chapter I -Mine Safety and Health Administration, 
Department of Labor; Subchapter - Coal Mine Health and Safety; 
Part 18 -Electric Motor-Driven Mine Equipment and Accessories; 
Part 75 -Mandatory Safety Standards, Underground Coal Mines; 
Part 77 -Mandatory Safety Standards, Surface Coal Mines and 
Surface Work Areas of Underground Coal Mines; 1981. 



20 



CHAPTER 2.— ELECTRICAL FUNDAMENTALS I 



The technique used to solve problems in complex elec- 
tronic circuits or mine power systems is called circuit an- 
alysis. It involves calculating such circuit properties as cur- 
rents, voltages, resistances, inductances, and impedances. 
Circuit analysis serves as the knowledge base on which an 
understanding of mine electrical systems can be built. 

This chapter will diverge from classical circuit-analysis 
presentations by not covering transient effects in circuits. 
From experience, studying currents and voltages existing 
in a circuit immediately after a change in circuit configur- 
ation can be confusing and clouds understanding of the most 
used segments of circuit analysis. Therefore, although some 
necessary statements will be made, the subject of transi- 
ents is delayed until chapter 11, where it can be combined 
with practical examples. 

This chapter commences by introducing electrical phe- 
nomena and continues through to a presentation of steady- 
state ac circuit analysis. Chapter 3, "Electrical Fundamen- 
tals II," continues the coverage of basic electrical subjects 
and starts with the basics of electrical power consumption. 

Numerous excellent circuit-analysis textbooks have 
been produced over the years. Many can be employed ef- 
fectively to cover the subject, and some of these are provided 
in the bibliography at the end of this book. Because practic- 
ally all fundamental electrical relationships are considered 
common knowledge, the concepts introduced in this chapter 
will seldom be referenced other than by giving credit to the 
discoverer. 

BASIC ELECTRICAL PHENOMENA 

The nature of electricity is not yet fully understood, but 
it is well known as a form of energy that can be conven- 
iently converted into and utilized as light, heat, and me- 
chanical power. Like all science, knowledge about electricity 
has been developed from observation and experimentation. 
The generalization of this experimental evidence combined 
with information about the nature and behavior of electrons 
and electron flow forms the basis of electron theory. 

The atoms of each element consist of a dense nucleus 
around which electrons travel in well-defined orbits or 
shells. The subatomic particles, the building blocks out of 
which atoms are constructed, are of three different kinds: 
the negatively charged electron, the positively charged pro- 
ton, and the neutral neutron. The negative charge of the 
electron, e~, is of the same magnitude as the positive charge 
of the proton, e*. No charges of smaller magnitude have yet 
been concretely observed. Thus the charge of a proton or 
an electron is taken as the ultimate natural unit of charge. 
It is these two particles that are of principal interest in 
electricity. 

Coulomb's Law 

The force, F, between two charges, q and q', varies 
directly as the magnitude of each charge and inversely as 
the square of the distance (r) between them. This relation- 
ship, known as Coulomb's law, is represented mathemat- 
ically by 



F = k^-, 
r 2 



If force is in newtons, charge in coulombs, and distance in 
meters, then 

k = 9 x 10 +9 N-nrVC 2 . 

The unit of charge, the coulomb (C), can be defined as the 
quantity of charge that, when placed 1 m from an equal and 
similar charge, repels it with a force of 9 x 10 +9 newtons 
(N). The charge carried by an electron or by a proton is 
e = 1.602 x 10- 19 C. 

Voltage and Current 

A proton in the nucleus of an atom can hold only one 
electron in orbit around it. When an atom contains fewer 
than the normal number of electrons that the protons can 
attract, the atom has an excess of positive charge and is 
said to be positively charged. Atoms with an excess of elec- 
trons are said to be negatively charged. The net amount of 
these charges is termed potential or electromotive force (emf) 
and is measured in volts. The separation of opposite charges 
of electricity may be forced by physical motion or may be 
initiated or complemented by thermal, chemical, or mag- 
netic causes or even by radiation. 

The potential difference or voltage existing between two 
points can be measured by the work necessary to transfer 
a unit charge from one point to the other. The volt is the 
potential between two points when 1 joule (J) of work is re- 
quired to transfer 1 coulomb (C) of charge. In other words, 

1 V = 1 J/C. 

In some metals or conductors, electrons in the outermost 
orbit of the atoms are rather loosely bound to their respec- 
tive nuclei. These are called conduction electrons, since they 
can leave the atom upon the application of a small force 
and become free to move from one atom to another within 
the material. In some materials, however, all the electrons 
are tightly bound to their respective atoms. These are called 
insulators, and in these materials it is exceedingly difficult, 
if not impossible, to free any electrons. Conductors and in- 
sulators are the principal materials used in electrical 
systems. 

The application of a voltage across a conductor causes 
the free electrons within the conductor to move. Electrical 
current is defined as the motion of electrical charge. If the 
charge in the conductor is being moved at the uniform rate 
of 1 coulomb per second (C/S), then the constant current 
existing in that conductor is 1 ampere (A), the unit of elec- 
trical current. The amount of current in a conductor can 
also be measured as the rate of change of the charge flow. 
Such changing current at any point in time is called in- 
stantaneous current or 



(the rate of change of charge) = -rf, 



(2.2) 



where i 



(2.1) 



where k = proportionality constant that depends on units 
used for force, charge, and distance. 



instantaneous current, A, 
q = flow of charge, C, 
t = time, s. 

When electricity was first discovered, it was erroneously 
thought that it was the flow of positive charges. Since the 
laws of attraction and repulsion were known, the movement 
was assumed to be from positive to negative. This theory 



21 



was accepted until the discovery of the radio tube, when 
it was recognized that the flow was movement of electrons 
from negative to positive. However, the concept of positive- 
charge flow was firmly entrenched and has remained stand- 
ard in the United States, and so it will be used here. 

SYSTEM OF UNITS 

Most material contained in this text is given in the 
International System of Units (SI); exceptions are calcula- 
tions that are more conveniently expressed in terms of the 
English or American engineering systems. A listing of the 
basic symbols, units, and abbreviations that are used is 
given in table 2.1. The decimal system is used to relate 
larger and smaller units to basic units, and standard pre- 
fixes are given to signify the various powers of 10; for 
example: 

pico- (p-, 10" 12 ) 

nano- (n-, 10~ 9 ) 

micro- (^-, 10" 6 ) 

milli- (m-, lO' 3 ) 

kilo- (k-, 10 3 ) 

mega- (M-, 10 6 ) 

giga- (G-, 10 12 ) 

Voltage, current, and power variables are represented 
by the letter symbols V, I, and P in both uppercase and 
lowercase letters. Uppercase letters represent voltage, cur- 
rent, and power when the variable is constant, as in dc 
circuits. In ac circuit work, uppercase V and I represent 
effective values and uppercase P represents average power. 
Lowercase v, i, and p depict voltage, current, and power 
when these quantities are varying with time. 



Where needed, double-subscript notation is used to de- 
scribe current and voltage. V AB represents the voltage of 
point A with respect to point B. I CD represents the current 
flowing through a circuit element from C to D. Note that 
in the circuit shown in figure 2.1, the voltage V^b causes 
the current l AB to flow. These meet with the standard for 
electrical current, which is positive-charge flow from posi- 
tive to negative. 



EXPERIMENTAL LAWS AND PARAMETERS 

It is remarkable that the entire theory of electrical 
circuits is based on only six fundamental concepts. One is 
Ohm's law, two are named for Kirchhoff, two relate to in- 
ductance and capacitance, and one has to do with power. 
To understand any electrical system, comprehension of 
these relationships is mandatory. 

Ohm's Law 

Georg Simon Ohm (1789-1854) discovered that the elec- 
trical current through most conductors is proportional to 
the voltage (potential) applied across the conductors. This 
phenomenon is known as Ohm's law and is expressed 
mathematically as 



= Ri, 



(2.3) 



where v = applied potential, V, 

i = current through the conductor, A, 
R = proportionality constant known as resistance 
of conductor, Q. 



Table 2.1— SI symbols and units 



Quantity 



Variable 
symbol 1 



SI 

unit 



Unit 


Identical 


symbol 


unit 


C 


As 


A 




V 


W/A 


V 




V 




Q 


V/A 


S 


A/V 





V/A 


S 


A/V 


Q 


V/A 


S 


A/V 


F 


C/V 


H 


Wb/A 


J 


N-m 


W 


J/s 


VA 




var 




Q-m 




S/m 




C 




C/m 2 




V/m 




F/m 




Wb 


Vs 


A/Wb 




H- 1 




Wb/A 




H 

T 


Wb/m 2 


A/m 




H/m 





Charge 

Current 

Voltage 

Electromotive force . 
Potential difference . 

Resistance 

Conductance 

Reactance 

Susceptance 

Impedance 

Admittance 

Capacitance 

Inductance 



Energy, work 

Power (active) 

Power — apparent . 
Power — reactive . . . 



Resistivity 

Conductivity 

Electric flux 

Electric flux density, displacement . 

Electric field strength 

Permittivity 

Relative permittivity 



Magnetic flux 

Magnetomotive force . 

Reluctance 



Permeance 

Magnetic flux density.... 
Magnetic field strength . 
Permeability (absolute) . 
Relative permeability.... 



I 

V.E...U 
V 
V,* 

R 
G 
X 

B 
Z 
Y 
C 

L 

W 
P 

S...P„ 
Q...P q 

p 

* 

D 

E 



* 

F...F 

R...R 

P...P 

B 

H 

V- 
Mr 



coulomb . 
ampere- 
volt 

...do 

...do 



ohm 

Siemens. 

ohm 

Siemens. 

ohm 

Siemens. 

farad 

henry 



joule 

watt , 

voltampere. 
var , 



ohm-meter 

Siemens per meter 

coulomb 

coulomb per square meter. 

volt per meter 

farad per meter 

(numeric) 

weber , 

ampere (amp turn) 

i ampere per weber 
reciprocal henry 
weber per ampere 
henry , 
tesla , 

ampere per meter 

henry per meter 

(numeric) 



V,E indicates alternative symbols;. ..U indicates reserve symbols. 



22 



No restriction is placed on the form of v and i. In dc cir- 
cuits they are constant with respect to time, and in ac 
circuits they are sinusoidal. 

For metals and most other conductors, R is constant. 
In other words, its value is not dependent on the amount 
of current, i. In some materials, especially in crystalline 
materials called semiconductors, R is not constant, and this 
characteristic is useful in diodes, amplifiers, surge arresters, 
and other devices. 

Further experiments by Ohm indicated that the resist- 
ance of a piece of metal depends on its size and shape. How- 
ever, the resistivity, q, of the metal depends only on its 
composition and physical state. This is an inherent prop- 
erty that opposes current through the conductor just as the 
frictional resistance of a pipe opposes the flow of water 
through it. Resistivity is defined as the resistance of a unit 
cube of homogeneous material; hence, resistivity can be 
thought of as a property of the material at a point. Its value 
remains the same at all points in a homogeneous conduc- 
tor, but if the material is not homogeneous, its resistivity 
can vary from point to point. The value may also vary 
greatly for different conductors. The concept of resistivity 
is often used in the grounding and distribution aspects of 
mine electrical systems. 

Using the definition, practical resistivity units would 
be ohm-centimeter (Q-cm) and ohm-inch (S2-in). However, 
resistivity is usually expressed in ohm-meter (Q-m) (SI) and 
ohm-circular-mill-foot (English). The ohm-meter is the re- 
sistance of a material 1 mm 2 in cross section with 1 m 
length. Likewise, the ohm-circular-mill-foot (usually abbre- 
viated to Q-cmil-ft) refers to the conductor resistance for a 
volume 0.001 in (1 mil) in diameter and 1 ft long. For 
calculating the resistance in this latter case, the cross- 
sectional area of the conductor is measured in circular-mills, 
which can be found from 



A = d 2 , 



(2.4) 



where A = cross-sectional area of circular conductor, cmil, 
and d = conductor diameter, 10" 3 in. 

Resistivity values of some common conductors are given in 
table 2.2. 

Table 2.2— Resistivity of some common materials at 20° C 

Material 

Aluminum, commercial 

Copper, annealed 

Iron, annealed 

Lead 

Nichrome 

Silver 

Steel, mild 

Tin 

Tungsten 



Temperature 


Resistivity (p) 


coefficient (a) 


10- 8 Q-m 


Q-cmil-ft 


0.0039 


2.824 


17.1 


.00393 


1.724 


10.5 


.005 


9.50 


57.4 


.0034 


21.83 


132.31 


.04 


100 


606.1 


.0038 


1.63 


9.85 


.002 


11.91 


72.17 


.0042 


11.50 


69.7 


.0045 


5.50 


33.2 



If I is in meters and A is in meters squared, then q must 
be given in units of ohm-meters. 

Electrical resistivity does not remain constant if the 
temperature is permitted to change. For most materials, 
the resistance increases as the temperature increases; car- 
bon is an exception to this rule (negative temperature co- 
efficient, 0.005). If the temperature coefficient is known, the 
resistance of a given conductor at a given temperature is 



R = R [1 + a (t - to)], 



(2.6) 



where R = resistance at temperature t, Q, 

Ro = resistance at reference temperature t , usually 

20° C, Q, 
a = temperature coefficient, Q/°C, 
t = given conductor temperature, °C, 
t = reference temperature, °C. 

At very low temperatures (about - 200 ° C for copper) or as 
the melting point is reached, the temperature coefficient 
is no longer constant and changes with temperature. As 
a result, equation 2.6 is not valid for very high or low 
temperatures. 

The symbol illustrated in figure 2.1 portrays a resistor 
in a circuit, and often its resistance is stated. Again by 
definition, 



R 



(2.7) 



Sometimes, the element's conductance, G, is referenced 
and is defined as the reciprocal of resistance: 



G = i = l. 
R v 



(2.8) 



In circuit analysis, it is occasionally more convenient to use 
conductance than resistance. Later, the explanation of this 
symbol will be generalized. 

Kirchhoff's Voltage Law 

In the simple series circuit shown in figure 2.2, three 
resistors are connected in tandem to form one single closed 
loop. Kirchhoff has shown that when several elements are 



r^WV 




Figure 2.1. - Circuit element Illustrating voltage polarity and 
current flow direction. 



The resistance of any specific conductor can be calcu- 
lated from the material resistivity using the formula 



R 



where R = resistance, Q, 

I = conductor length, 

A = conductor cross-sectional area, 
and q = material resistivity. 



(2.5) 




Figure 2.2. — Simple series circuit. 



23 



connected in series, the current in the circuit will adjust 
itself until the sum of voltage drops in the circuit is equal 
to the sum of voltage sources in the circuit. This can be 
restated as the "sum of all voltages around any closed cir- 
cuit is zero," which is called Kirchhoff s voltage law. For 
the circuit shown in figure 2.2, 



or 
or 



v oi + v ic + v cd + v da = 

v afc + v bc + v cd - v ad = 

v 1 + v 2 + v 3 - v s = 0. 



(2.9) 
(2.10) 
(2.11) 



Obviously, some of these potential differences could be 
negative and some positive. This circuit shows only resist- 
ances and a voltage source, but the network could contain 
other kinds of elements and might be as complicated as 
desired. However, Kirchhoff found that the sum of the volt- 
ages around any closed loop in a circuit, such as a-b-c-d, is 
always zero. 

The symbol shown in figure 2.2 beside v s represents an 
ideal voltage source. Such a source maintains a given 
voltage across its output (terminals) regardless of the load, 
but actual voltage sources cannot supply an infinite cur- 
rent if the terminals are short-circuited; that is, they are 
tied together so the resistance approaches zero. Therefore, 
actual sources are usually considered to be ideal voltage 
sources with an internal resistance connected in series with 
the source and the output terminals. The assumption is 
illustrated in figure 2.3. 



EXAMPLE 2.1 

Find the current I flowing in the single-loop cir- 
cuit in figure 2.4. 

SOLUTION. Adhering to the assigned clockwise 
direction for current, Kirchhoff s voltage law produces 
the following equation: 

-50 + V, + 100 +V 2 = 0, 

where V, and V 2 are the voltages across the l-£2 
and 2— Q resistances. From Ohm's law, 

V, = II, 
V 2 = 21. 

Inserting these expressions into the voltage law equa- 
tion produces 



or 



-50 + II + 100 + 21 = 
31 = -50, 
I = -16.7 A. 



The negative sign states that the actual current flow 
is in the opposite direction from that shown in fig- 
ure 2.4. 

It should be noted that when writing the voltage- 
law equation, voltages that oppose the assigned cur- 
rent flow are considered positive, otherwise negative. 
Therefore, the 100-V source is positive, and the 50-V 
source is negative. The positive signs for V! and V 2 
assumed opposition by the convention shown in fig- 
ure 2.1. 



Kirchhoff's Current Law 

The other law attributed to Kirchhoff specifies that 
"the sum of all electrical currents flowing toward a junc- 
tion is zero." In figure 2.5, five wires are soldered together 
at a common terminal and the current in each wire is 
measured. If current flowing toward the junction is called 
positive (the direction shown in the figure) and the current 
outwards is negative (against the arrows), then the sum of 
the five currents is zero: 



i, + i, + i, + L + i. = 0. 



(2.12) 



As was the case for equation 2.9, this equation implies that 
some currents must be positive, some negative. 

If two or more loads are connected between two com- 
mon points or junctions, these elements are said to be in 
parallel, as shown in figure 2.6A. The same is true for figure 
2.6B, and moreover, the two circuits illustrated in figure 
2.6 are identical, just drawn differently. It is important to 




t/W- 



I O Resistance 

A 

Ideal — — 

source 



Output 
terminals 



,X 



Ideal voltage source Actual voltage source 

Figure 2.3. — Ideal and actual voltage sources. 




Figure 2.4. -Circuit for example 2.1. 




Figure 2.5. — Demonstration of Kirchhoff's current law. 



24 



note that the lines in these and all circuit diagrams usu- 
ally show no resistance. Each line is only a connection be- 
tween elements or between an element and a junction. The 
similarity in the diagrams can be shown using KirchhofFs 
current law. In both, there are only two independent junc- 
tions, a and b, and for either point, 



ii = i 2 + i 3 + U 



(2.13) 



The circuit symbol next to i x in figure 2.6 represents 
an ideal current source, and a similar situation exists for 
all practical current sources as was mentioned for practical 
voltage sources. However, the internal resistance is effec- 
tively connected in parallel across the ideal current source. 
Both ideal and actual current sources are shown in fig- 
ure 2.7. 



EXAMPLE 2.2 


Verify that KirchhofFs current law holds for junc- 
tion x in figure 2.8. 


SOLUTION. The three resistances in figure 2.8 are 
in parallel, and the 100 V produced by the voltage 
source exists across each. Therefore, by Ohm's law, 
the current through each resistance is 


T = 100 = 4 a 

■•■2S * ■**! 


25 


I 50 = 100 = 2 A, 


50 


I = 100 - i a 
i 100 — -■-"" — ± t\. 


100 


KirchhofFs current law states that for junction x, 


I25 + I50 + lioo = 7 A. 


Accordingly, 

4 + 2 + 1 + = 7 A. 



Series Circuits 

To restate the earlier definition of a series circuit, 
elements are said to be connected in series if the same cur- 
rent passes through them. Such is the situation for the four 
resistors shown in figure 2.9. It would be convenient to find 
a resistance, R, that could replace all series resistors. This 
equivalent resistance can be found by returning to the Ohm 
and Kirchhoff voltage laws. By Kirchoff s law, 



V = v, + v 2 -I- v 3 + v„, 
but by Ohm's law, 

v, = iRj, v 2 = iR 2 , v 3 = iR 3 , - 
Therefore, 



For the circuit in figure 2.9, if the voltage, v, produces the 
same current, i, through the circuit, then 



but 



or 



v = iR, 
v = i (R x + R 2 + R 3 + R 4 ), 
iR = i (R, + R 2 + R 3 + R 4 ), 
R = R x + R 2 + R 3 4- R 4 . 



(2.14) 



Here R is said to be the equivalent resistance for the 
previous series circuit. In other words, R is the series 
resistance of that circuit. The same logic applies to all elec- 
trical elements in series. 




a 

* •- 



Q) R. 



»" i\" ^ 



A B 

Figure 2.6. -Simple parallel circuits. (Italic letters are cited 
in text.) 



Ideal 

current 

source 



1 . 


Ideal S\ 


5!*"* Resistance 




current 
source 





Output 
terminals 



Actual current source 
Figure 2.7. — Ideal and actual current sources. 



IOOV 




25/1 <>5on. <ioo/i 



Figure 2.8. — Parallel circuit for example 2.2. 



R. 



^AAAr 
R, 



jyv\/\r 



^V\AAr 



v = iR! + iRj + iR 3 + iR 4 . 



Figure 2.9.— Simple series circuit and equivalent. 



25 



It is often useful to find the voltage drop across just one 
element in a series circuit. To arrive at an expression, again 
refer to figure 2.9. For the current through the circuit, it 
is obvious that 



1 = i x = i 2 = i s = i 4 



but 
Therefore, 



ii = 



Ri* 



R,' 



Rj R 2 R 3 R 4 



As before, consider R, the equivalent circuit resistance, and 

v 



Therefore, 



X ~R- 

v. _ Zl v. _ v* 
R Rj R R 2 



or 



5l 
R 



R 



(2.15) 



In other words, the voltage drop across any one element is 
equal to the total circuit voltage times the ratio of the ele- 
ment's resistance to the total circuit resistance. 

Parallel Circuits 

Following the discussion of series circuits, it would be 
useful to have similar equivalence, voltage, and current 
relationships for parallel circuits. For the circuit shown in 
figure 2.10, the voltage is the same across each resistor and 
is a corollary to current through series elements. Using the 
same basic procedures as for series circuits, it can be shown 
that 



G = Gi + G 2 + G 3 + G 4 + 



and also 



-^ = ^. + i + i + i + 
R Rj R2 R3 R 4 



(2.16a) 



(2.166) 



Restated, the total conductance of parallel-connected re- 
sistors is equal to the sum of all individual conductances. 
Likewise, the reciprocal of the total resistance of parallel- 
conducted resistors is equal to the sum of the reciprocals 
of the individual resistances. A special case that is very 
often found occurs when two resistances are in parallel. If 
these resistances are Ri and Rj, then 



jj _ R1R2 



Ri + R2 



(2.17) 



If current distribution through parallel circuits is of in- 
terest rather than voltage distribution, Kirchhoff s current 
law and Ohm's law can be employed to show that 

Gi R 

i 1= -fiori^ f-i, (2.18) 

and so on for the balance of currents. 

The preceding paragraphs have been used to show the 
immediate application of Ohm's law plus Kirchhoff s volt- 
age and current laws to circuits that have more than one 
element. The results are extremely valuable in circuit anal- 
ysis and are used extensively to solve circuit problems. It 
is important to note now that these concepts are also valid 
when circuits contain components other than resistance. 



Later, after the balance of fundamental laws and param- 
eters have been covered, more applications of these laws 
will be shown. 



EXAMPLE 2.3 

A series-parallel circuit is shown in figure 2.11. 
Find the equivalent resistance. 

SOLUTION. The objective is to find an equivalent 
resistance between terminals a and b. The process is 
to combine resistances in series or in parallel until 
the equivalent resistance is obtained. The 2-Q and 4-Q 
resistances between point 1 and terminal b are in 
series, and from equation 2.14, 

2 + 4 = 62. 

If a 6-Q resistance replaces these two series resist- 
ances, it can be seen that two 6-Q resistances are in 
parallel between point 1 and terminal b. Applying 
equation 2.17, 

(6X6) 



6 + 6 



3 Q. 



which means that a 3-Q resistance can replace the two 
6-Q parallel resistances. Therefore, the 3-Q resistance 
between point 2 and point 1 is in series with the 
equivalent of 3 Q between point 1 and terminal b, and 
again 

3 + 3 = 6 Q. 

Now between point 2 and terminal b, there are the 
equivalent of two 6-Q resistances in parallel and 

6 + 6 

Consequently between terminal a and terminal b, a 
7-Q resistance is in series with an equivalent 3-Q 
resistance, and the equivalent resistance of the en- 
tire circuit is 

R = 7 + 3 = 10 Q. 



: R 1 < R 2 ^ R 3 



Figure 2.10. — Simple parallel circuit. 



7A 2 3n 1 2A 
a o — v\A/ — •— <wv- 




Flgure 2.11. — Series-parallel circuit for example 2.3. 



26 



EXAMPLE 2.4 

Find the equivalent resistance of the circuit illus- 
trated in figure 2.12. 

SOLUTION. Point b and point b' can be seen in the 
center of the circuit, but these are electrically just one 
point, because b and b' are only separated by a line 
that does not contain an electrical element. Thus, the 
15-Q and 30-Q resistances between a and b are in 
parallel, as are the two 40-Q resistances between b 
and c. 
From equation 2.17, 



and 



(15X30) 
15 + 30 

(40X40) 
40 + 40 " 



= 10 Q 



20 Q 



Therefore, the resistance of the circuit between ter- 
minal a and terminal d can be reduced to three series 
resistances, and the equivalent resistance is 

R = 10 + 20 + 10 = 40 Q. 



The Magnetic Field 

A. M. Ampere was the first scientist to establish that 
the conductor through which electric current is passing is 
enclosed in a magnetic field. The relationship is depicted 
in figure 2.13A. After Ampere's discovery, many experimen- 
ters tried to reverse the process and create electric current 
from a magnetic field. Finally, in 1831, Michael Faraday 
discovered that as a magnet is inserted into a coil of wire, 
an impulse of electrical current will flow through the wire. 
When the magnet remains stationary within the coil, no 
current is produced. When the magnet is withdrawn, a cur- 
rent impulse is again observed, but this time it flows in the 
opposite direction. The process is demonstrated in figure 
2.14. Faraday visualized the effect as a result of magnetic 
flux lines cutting or moving through the conductor. When- 
ever relative motion occurs, an emf is produced in the 
conductor. This disclosure laid the foundation for electro- 
mechanical conversion, that is, the conversion from mechan- 
ical energy to electrical energy and vice versa, as found in 
generators and motors. 

The magnetic field mentioned here is a condition of 
space. The direction of a magnetic field flux line is the direc- 
tion of force on a magnetic pole, and the flux line density 
is in proportion to the magnitude of force on the pole. Each 
line represents a certain quantity of magnetic flux, meas- 
ured in webers. It is a magnetic field characteristic that 
every flux line is a closed curve, forming the concentric- 
circle pattern shown in figure 2.13A. These conditions 
of the magnetic field are employed to develop relation- 
ships in magnetic devices, which are covered in upcoming 
sections. 

When a wire is wound into a coil, an interesting action 
occurs: as the magnetic flux builds up around one wire, it 
tends to cut through adjacent turns of wire. In this way a 
voltage is induced into the coil windings. The concept is 
shown as dashed flux lines around one winding of the coil 
in figure 2.13B. 



30A 




40X1. 



do M/V*- 

ion c 

Figure 2.12. -Series-parallel circuit for example 2.4. 




Lines of 
magnetic flux 



Magnetic flux 




Current 



B 



Figure 2.13. — Magnetic flux In a straight conductor (A) and 
in a long coll (B). 



Cardboard 
torm CoH 



Bar magnet moving 
into the coil 



Magnet moving out 
of the coil 





Galvanometer Current 

Figure 2.1 4. -Demonstration of induced current. 

Inductance 

Joseph Henry found that electricity flowing in a circuit 
has a property analogous to mechanical momentum; that 
is, current is difficult to start but once started it tends to 
continue. This is the case for any element from a simple 
conductor to the most complex. Faraday explained the phe- 
nomenon by visualizing the magnetic field in space around 
the conductor. In terms of the coil in figure 2.14, the volt- 
age induced in the other windings is proportional to the rate 
at which the magnetic flux lines are cutting through the 
coil. Yet the magnetic flux is also proportional to the cur- 
rent in the coil. The induced voltage is such that at every 
instant, it opposes any change in the circuit current. For 



27 



this reason, the induced voltage is called a counterelectro- 
motive force, cemf. This interrelationship is so important 
that it has the status of a physical law and is known as 
Lenz's law after the scientist who first defined it. 

The property that prevents any change of current in the 
coil is called self-inductance; hence the coil is known as an 
inductor. The greater the induced voltage, the greater is 
the opposition to the change in current flow. Therefore, the 
cemf produced by a specific change of current is a measure 
of circuit inductance. Expressed as a formula, 

v = IXrate of change of current) 



or 



v = L( dT x 



(2.19) 



where v = voltage across coil, V, 

L = proportionality constant known as inductance, H, 
i = current through coil, A. 

As noted, inductance is given the symbol L and is measured 
in units called henries in honor of Joseph Henry. A circuit 
has an inductance if 1 H when a current change of 1 A/S 
causes a cemf of 1 V to be induced in the coil. The expres- 
sion "di/dt" represents the rate of change of current, i, in 
the coil. 

When two separate coils are placed near each other, as 
shown in figure 2.15, the magnetic field from one coil can 
cut through the windings of the second coil. It follows that 
a change in the current in coil 1 can produce an induced 
voltage in coil 2. This current-voltage relationship is ex- 
pressed as 

v 2 = L 21 (rate of change in i t ) 



or 



dii 

L 21 (— ). 
1 dt 



(2.20) 



here to demonstrate the parameters that affect inductance. 
The two symbols used to indicate inductance are shown in 
figure 2.16; the symbol on the right is that commonly found 
in power-circuit diagrams. 

For a long coil as shown in figure 2.16, the inductance is 



L = 



M N 2 A 



(2.23) 



where L = self-inductance, H, 

\jl = permeability, H/m 

(for air, 4n-10" 7 = 12.56610- 7 ), 

N = turns of coil, 

A = coil cross-sectional area, m 2 , 
and ( = coil length, m. 

The coil cross section need not be circular. The formula is 
only approximate because it assumes that all flux lines link 
all turns of the coil, which cannot occur at the coil ends. 
However, the formula gives good results for long coils and 
does reveal the following important relationships. 

• Coil inductance is proportional to the square of the 
number of turns. 

• Inductance is proportional to the core permeability. 

• Inductance is proportional to the cross-sectional area 
of the core. 

• Inductance decreases as the length increases. 

For a shorter single-layer circular solenoid (coil), the induc- 
tance is approximately 



L = /iN 2 A 
I + 0.45d' 



(2.24) 



where d = coil diameter, wire center to center, m. 



Similarly, if the current in coil 2 is changing, it induces a 
voltage in coil 1: 

v i = Li 2 (rate of change in i 2 ) 



or 



at 



(2.21) 



L 12 and L 21 are called mutual inductances and are again 
expressed in henries. The mutual inductances increase if 
the coils are brought closer together and decrease as the 
coils are moved further apart. Two magnetically coupled 
coils are usually called a transformer. Although not by all 
means obvious, the two mutual inductances of a pair of 
magnetically coupled circuits are equal, or 



L 12 — L, 



(2.22) 



The self-inductance of an actual coil is a function of both 
the coil configuration and the total number of turns. Fur- 
ther, because the magnetic flux may induce currents in ad- 
jacent conductors, the environment in which the coils are 
placed may also have an effect. Numerous inductance equa- 
tions are available in handbooks and other reference books, 
each valid for a given coil configuration; consequently, only 
a few that give approximate inductance values are provided 



12 
L 2 i 



'2 



CoiM 



Coil 2 



Figure 2.1 5. -Two colls demonstrating mutual Inductance. 



Symbol 

o— ^mnnr^-o 



d \*—r 
1 ' I 



Area 
A 



\y ^^^m 



It 



N turns 



Symbol 



Figure 2.16. -Long-coll Inductance and Inductor symbols. 



28 



For the toroidal coil of rectangular cross section in figure 

2.17, 

L-^pliA (2.25) 

2n d t 

where d„ d, = inner and outer diameters as shown, m, 
and h = thickness, m. 

Note that In indicates the natural logarithm, that is, the 
logarithm to the base e. 

Capacitance 

When two conducting surfaces are separated by a dielec- 
tric or insulating material, an effect known as capacitance 
is observed. If two electrical conductors are at different 
potential, there is some storage of charge upon them. A 
capacitor is a device included in a circuit for the purpose 
of storing or exchanging this electrical charge. Further, 
when capacitance is present, the charge observed to flow 
into the capacitor is proportional to the voltage applied. This 
can be expressed as: 



Cv, 



(2.26) 



where 



and 



q = stored charge in capacitor, C, 

v = applied voltage, V, 

C = proportionality constant called capacitance, F. 



To analyze circuits, a relationship between the voltage 
applied and the current flowing into and from the capaci- 
tor is more useful. Current is the rate at which charge flows 
(i = dq/dt). It therefore holds that for a given capacitance, 

i = C(rate of change of v) 

= (A, (2.27) 

dt 

where i = current flowing into the capacitor, A, 
and v = voltage across capacitor, V. 

This is very similar to equation 2.19 and, using the discus- 
sion in that section, capacitance can be defined as that elec- 
trical circuit property which tends to oppose any change in 
voltage. The capacitance of a capacitor depends on the size 
of the conductors or plates, their proximity, and the nature 
of the material between them. For most dielectric materials, 
C is constant. 

Equations 2.26 and 2.27 have algebraic signs consistent 
with the arrows in figure 2.18. The symbol shown is for 
capacitance; note that a positive terminal voltage produces 
positive current and hence positive charge. 

If the voltage across a capacitor is desired, equation 2.27 
can be integrated, resulting in 



v = ±f m idt + V.. 



(2.28) 



This equation represents the change in voltage across the 
capacitor from some arbitrary reference time, called t = 0, 
to a later time, t. V is the potential across the capacitor 
at time t = 0. The expression 

"Trf idt" 
C • 

is the voltage change across the capacitance from time 

t = to time t = t. From the formula, if the voltage across 



the capacitor remains constant, as in dc circuits, no current 
will flow into or out of the capacitor. 

Electric Field 

An electric field exists anywhere in the neighborhood 
of an electrical charge, for example, between the plates of 
a capacitor. The direction of this field is by definition the 
direction of the force on a positively charged exploring par- 
ticle (a particle free to move within the electric field). The 
strength of the field, E, is proportional to the magnitude 
of the force. If the charge of the exploring particle is q, then 
the force is 

F = qE, (2.29) 

where F = force on particle, N, 

and E = strength of electric field, V/m. 

Electric-field flux lines are visualized as issuing from 
positive electric charge and terminating on negative charge 
as shown in figure 2.19. 







IE- 



Cross 
section 



N turns 



Figure 2.17. -Toroidal coil. 



o— 


i - 


— 




+ ' 


1 






V 


+ + 


+ + 










— — 


— - 


o— 




q 





Symbol 

1 

T 



Figure 2.18. -Charge, voltage, and current relationships of 
capacitor. 




Figure 2.19. -Electric lines of force between two parallel 
charged plates. 



29 



Voltage or potential difference is by definition the in- 
tegral of electric-field strength or 



JE-ds. 



(2.30) 



A simple application of this concept can be demonstrated 
from figure 2.19. If the electric field between the two parallel 
plates is constant, the voltage between the plates is 



= E-s. 



(2.31) 



Assume that a positively charged particle, q, is released 
from the positive plate in figure 2.19, the particle being 
within the electric field and free to move. If it moves, 
work is performed on it by the electric field. The amount 
of work can be found by employing the mechanical formula 



w = Fs, 



(2.32) 



where 

and 

Since 
work is 
but 
so 



w = work done, J, 

F = force on particle, N, 

s = distance particle moves, m. 

F = qE, 
w = qE-s, 

v = E-s, 
w = qv. 



(2.29) 
(2.33) 
(2.31) 
(2.34) 



Therefore, when electricity moves from one potential to 
another, the work done is equal to the product of the amount 
of electricity and the potential difference. In the next 
section, this concept is applied to a common electrical 
component. 

Instantaneous Power 

Consider the resistor shown in figure 2.20. A charge, 
dq, is free to move in the resistor from the point s to s + ds. 
It moves the distance, ds, in time, dt, and is impelled by 
the electric field in the region, E. 

The electric field exerts a force on the charge, dq, while 
it moves through ds, or 



F = dqE. 



(2.35) 



The work done in this section of the resistor during time 
dt can be expressed as 



dw = F-ds = dqEds. 



(2.36a) 



+ A 




s + ds 



Figure 2.20. - Resistor used to demonstrate Instantaneous 
power. 



Power is work per unit time (in other words, the rate of do- 
ing work), or for this section of the resistor, 



dp 



where p = power, W. 



dw _ dqE-ds 



dt 



dt 



(2.366) 



However, the current through the resistor is the rate at 
which charge flows, i = dq/dt; therefore, 



dp = iE-ds. 



(2.37) 



Current is constant throughout the resistor and is not a 
function of distance, s. The potential difference across the 
region, ds, is 

v = / E-ds, (2.30) 

and the power across the whole resistor, from a to b, is then 

p = j'iE-ds =i jT*E-ds = iv. (2.38) 

Formula 2.38 represents only the instantaneous power con- 
sumed by the resistor, or the power occurring at only one 
instant in time. This is an extremely important formula as 
it forms the basis for most power relationships. 

Idealization and Concentration 

The foregoing has established the elementary laws and 
parameters that can be applied to investigate electrical 
circuits. Practical circuits found throughout a mine, or in 
fact anywhere else, are composed of wires, coils, and elec- 
trical devices of varying complexity. Before these funda- 
mentals can be employed, it is necessary to translate the 
practical world into an ideal and simple world. The trans- 
lation is called idealization and is in essence the construc- 
tion of a model. Here, electrical effects that create insig- 
nificant results are eliminated. For instance, two adjacent 
conductors in a coil always exhibit capacitance but the 
capacitance might be so small that the stored charge is 
negligible. Yet for many situations, the resistance and in- 
ductance must remain. 

For every conductor or component in a circuit, resist- 
ance, capacitance, and inductance are distributed through- 
out the entire length or breadth of the portion. It would be 
much simpler to apply the preceding relationships if these 
circuit parameters were combined or concentrated into 
separate circuit elements. For most circuit analysis needs, 
fortunately, these can indeed be consolidated. 

The fundamental aspects of idealization and concentra- 
tion are illustrated in figure 2.2LA, which shows a voltage 
generator connected to a coil of wire and a resistor in series. 
Figure 2.2 LB gives the translation. The distributed resist- 
ance and inductance of the coil have been combined into 
R £ and L, and the coil's capacitance has been ignored. The 
resistance of the resistor and its lead wires has been con- 
centrated into one value, R. Finally, the voltage generator 
is represented by an ideal voltage source and an internal 
series resistance. Note that the lines shown in figure 2.2LB 
serve only to connect components and exhibit no electrical 
properties or effects. Another example can be expressed 
from figure 2.22A. Here, a load center is shown connected 
to a shuttle car through a trailing cable. Again, figure 2.22B 
gives the translation. The distributed resistance, induc- 
tance, and capacitance of the trailing cable have been 



30 



Voltage 
generator 



Wire coil 



Wire ,. , Wire 
vVW 



Connection 



Resistor 
A 

Inductor 




Connection 



Connection 



Resistor 



Flgure 2.21 . — Simple example of Idealization and concentra- 
tion. 



Trailing 
cable 



Shuttle car 





A/VW- 



R tc 

J VWr 



-tc 



Motor 



uniform values. Beyond this, the term "dc" is also applied 
to ordinary or practical currents that are approximately 
steady. 

The following section explores dc circuit analysis, an 
important topic because of its extensive use for mine haul- 
age and for driving electronic components. The study of dc 
analysis at this time allows the fundamental electrical laws 
and parameters to be applied and extended without hav- 
ing the effort clouded by complex current relationships. 

Direct Current and Circuit Elements 

Figure 2.23 gives the basic elements of resistance, in- 
ductance, and capacitance, each having a voltage and cur- 
rent as shown. A powerful simplification of complex circuits 
can be understood by examining the effect of dc on these 
elements. As before, the voltage-current relationship for the 
resistor is Ohm's law: 

V„ = IR. (2.39) 

For the inductor, the voltage across the element is 

dl 

v < = L 7t' 

but, because I does not change, 
dl 



= 



and 



dt 



V, = 0. 



(2.40) 



Likewise, for the capacitor, the current through the element 
is 

CdV 



Ic = 



dt 



Figure 2.22. -Modeling of load center, trailing cable, and but again, 

shuttle car. 



represented by the combined R, c , L„, and C, c . Such models 
can be developed for all portions of a mine distribution 
system. In this case, although the cable capacitance is 
rather small, it is shown here to emphasize that it may not 
always be negligible. The shuttle-car motor is depicted by 
the symbol shown. The symbol is obviously the same as that 
for a source, for reasons given later. 

After constructing the circuit representation, or sche- 
matic, as it is most often called, the relationships covered 
previously in this chapter can be used to solve for currents 
and voltages within the circuit. A differential equation will 
usually result, but if the circuit contains only resistance, 
the only necessary expression is Ohm's law. This will be 
the case for most dc circuit analyses. 

DIRECT CURRENT CIRCUITS 

Electrical current consists of the motion of electrical 
charges in a definite direction. The direction and magni- 
tude of current can vary with time, and accordingly, all cur- 
rents can be classified into one of three basic types: 

Direct current (dc), 

Alternating current (ac) or sinusoidal current, and 

Time- varying current. 
Direct current is a steady, continuous, unidirectional flow or I 

of electricity. In other words, voltage (V) and current (I) have 



^ = 0, 
dt 



Ic= 0. 



(2.41) 



Therefore, inductance and capacitance phenomena are not 
present under pure dc. In other words, the capacitor appears 
as an open circuit, while an inductor resembles a conduc- 
tor showing only resistance. An example of this simplifi- 
cation is available in figure 2.24. The circuit on the left 
shows all circuit elements, but under dc the effective cir- 
cuit is given on the right. The result is a simple series- 
resistance arrangement, and the only voltage-current rela- 
tionship necessary for the analysis is again Ohm's law. 

Series and Parallel Resistance 

The expressions used to find the equivalent resistance 
of parallel or series resistances are as before: 



for series, R„, = R t + R2 + R 3 + .... + R„; 

for parallel, = -I- + + .... + 

R e , R x R 2 R 3 

using Ohm's law, V = IR,, = I(R L + RJ 

V r V 
R«, Rz. + Ri 



31 



Resistance 



Inductance 



Capacitance 



R 

L 

V 
c 



Ohms law: V = IR 
Power: P = VI 

V=l_§=0, Power = 



I = C^ = 0, Power=0 



Figure 2.23. — Basic elements of resistance, inductance, and 
capacitance. 



i — VWV- 



6 




Figure 2.24. -Simplification of dc circuit. 



Further, the current and voltage distributions can be deter- 
mined by 



L 



rt 2 



or 



R, 

v, = (— )V 



and 



and 



and 



L 



R 2 



Ro 

V 2 = (— )V, 

tXeq 

V 2 = R 2 'I t . 



In this way, all voltages and currents in figure 2.25 can be 
found. In summary, the main process used is the substitu- 
tion of a single resistance for several series-parallel resist- 
ances. In concept, the same terminal resistance (R e ,) implies 
equivalence and results in identical current and voltage 
delivered from the source. This process of solution is for- 
mally known as circuit reduction. 

The power consumed by all or part of the circuit can be 
found by applying equation 2.38: 



VI. 



(2.42) 



Noting Ohm's law, V = IR, two other convenient power ex- 
pressions are 

P = (IR)I = PR (2.43) 




.if] 



Ro' 



A B C 

Figure 2.25. - Simple circuit reduction. 

Hence, the current through or the voltage across the cir- 
cuit can be found. By employing the previously given for- 
mulas, the voltage-current relationships for each circuit part 
can be determined. 

This concept can be elaborated considering figure 2.25, 
which shows a series-parallel circuit where an element of 
the circuit may be in parallel or series with other elements. 
Arrangements such as these can be solved by observing 
which individual elements are in series or parallel, then 
making the appropriate combinations. The objective is to 
gradually reduce the circuit to an equivalent series arrange- 
ment, which can then be replaced by a single equivalent 
resistance. Simple illustrations of this process have been 
shown in examples 2.3 and 2.4. Accordingly, the circuit in 
figure 2.25A can be changed to figure 2.25JB by the paral- 
lel combination 



1 11 T, , 

■— ■ = 1 or R 2 

Rz' R2 R3 



R»Ra 

Ro + R-i 



The circuit in figure 2.25S is then reduced to the circuit 
in 2.25C for the series combination 



R. 



R 1 + Ro. — R 1 + 



R 2 Rs 
R, + R, 



Afterwards, if V is known, I x = 



and 



V V 2 
P = V(— ) = — . 
R R 



(2.44) 



These three expressions can be used to find the power loss, 
expressed as PR loss, due to conductor resistance before the 
current is delivered to a load, as well as the power used by 
that load. 



R. 



EXAMPLE 2.5 

For the circuit shown in figure 2.26, determine the 
current I flowing through the 30-Q resistance, the 
power supplied to the circuit by the voltage source, 
and the power consumed by the 15-Q resistance. 

SOLUTION. The 25-Q, 15-Q, and 10-Q series 
resistances are in parallel with the 50-Q resistance, 
and 



25 + 15 + 10 

(50) (50) _ 
50 + 50 



= 50 Q, 
25 Q. 



The equivalent resistance seen by the 50-V ideal 
source is the sum of three series resistances: 

R„ = 30 + 25 + 45 = 100 Q. 

The current delivered by the source is then 

V 50 



Ix = 



0.5 A, 



R«, 100 

and the power supplied to the circuit by the source is 
P, = VI, = 50(0.5) = 25 W. 



32 



The current through the 15-Q resistance can be found 
by using current division: 

50 50 

I 2 = U ) = 0.5( ) = 0.25 A. 

50 + 25 + 15 + 10 100 

Therefore, the power consumed by the 15-Q resistance 

is 

P = I1R = (0.25H5 = 0.94 W. 



EXAMPLE 2.6 

Find the current between points a and b, I a4 , in 
the circuit of figure 2.27. 

SOLUTION. The circuit is very similar to figure 2.12 
as used in example 2.4. Point a and point b are at the 
same potential, so the right-hand side of the circuit 
is essentially two parallel arrangements of 15-Q and 
30-Q resistances. The two parallel arrangements are 
in series. As 



(15) (30) 



10 Q, 



15 + 30 

the equivalent resistance seen by the 30-V source is 
R c , = 10 + 10 + 10 = 30 Q, 



and total circuit current is 
V _ 30 

1 ~ T " 30 



1 A. 



Because of effective parallelism, I a , I 6 , I c , and l d can 
be found by 

30 30 

I. = Ii( ) = K — ) = 0.67 A, 

15 + 30 45 



K- 



15 



Ic = K 



1<- 



15 + 30 

15 



15 + 30 

30 
15 + 30 



) = 0.33 A, 
) = 0.33 A, 
) = 0.67 A. 



Even though the potential at point a is the same as 
the potential at point b, the line connecting a to b has 
the ability to carry current. By Kirchhoff s current 
law for point a, 

I. = L* + Ic, 



and for point b, 



L + L = L- 



By either relationship, 

L„ = 0.67 - 0.33 = 0.33 A. 



EXAMPLE 2.7 

The circuit in figure 2.28 is a series-parallel ar- 
rangement of conductances. Find the voltage V. 

SOLUTION. As total conductance of parallel- 
connected conductances is the sum of the individual 
conductances, the combination of the two elements 
between points a and b is 

G ab = 2 + 2 = 4. 

This combined conductance is in series with the 1-S 
conductance so that total conductance of the circuit 
portion from point a through point b to point c is 

1 1 1 

Gabc G ab G bc 



or 



G a 



G a6 Gic 



(4X1) 

4+1 



0.8 S . 



G atc is in parallel with the 2-S conductance between 
points a and c, and the equivalent conductance seen 
by the ideal by the ideal current source is 

G„ = 2.0 + 0.8 = 2.8 S. 

Therefore, the voltage shown in figure 2.28 is 

I 28 



a 



2.8 



= 10 V. 



50 V 



-Ii»3on 25n ~Jz 

I WA f V\M 1 | 

6 f° n I 

I — vwv — • — vwv — I 



15X1 



45n ion 
Figure 2.26. - Circuit for example 2.5. 




30 V 



I jw 

ion 
Figure 2.27. - Circuit for example 2.6. 



2S 

1-WVW 1 

2S 

<>-t/W — ♦ 



28A (T 



V $2S §1S 



Figure 2.28. -Series-parallel conductances for example 2.7. 






33 



EXAMPLE 2.8 

For the circuit shown in figure 2.29, find the total 
circuit current, I a , with the components as shown, 
with the 5-Q resistor short-circuited, and with the 
5-Q resistor open-circuited. 

SOLUTION. For the circuit as shown, the two 10-Q 
resistors between points c and d are in parallel and 

_ mm _ 5a 

10 + 10 

This resistance is in series with the 5-Q resistance and 
these three elements are in parallel with both the 
15-Q and 10-Q resistances between points b and d. 
Thus, 

R.C = 5 + 5 = 10 Q 



and 



or 



R bd is in series with the 7-Q resistance and 

Raw = 7 + 3.75 = 10.75 Q, 

and R abd is in parallel with the 10-Q resistance be- 
tween points and a and d; both are across the 120-V 
source. Therefore, 



1 

Rid 


= 


_1 
15 


10 


1 
•Kicd 


1 
Rid 


= 


_1 

15 


+ -1+ 

10 


1 
10 




R 


>d = 


3.75 Q. 





R, 



(10.75) (10) 
10 + 10.75 



and total circuit current is 

V = _120 
R,„ " 5.18 



Ix = 



= 5.18 Q, 



23.2 A. 



When the 5-Q resistance in figure 2.29 is short- 
circuited or replaced with zero resistance, points b and 
c are electrically the same. Four resistances are now 
in parallel between points b and d; three 10-Q and the 
15-Q resistance. Following the same procedure as 
before, the equivalent resistance becomes 

K'„ = 4.93 Q, 

and total circuit current is 
120 



V 



4.93 



24.3 A. 



For the case of an open-circuited 5-Q resistance, the 
resistance between b and c is assumed to be infinite, 
and the two 10-Q resistances between points c and d 



are disconnected from the circuit. The equivalent 
resistance is now 

R4' = 5.65 Q, 

and the total circuit current is 
120 



Ix" 



5.65 



21.2 A. 



This example illustrates an important concept. 
When an element in a circuit is short-circuited, the 
equivalent resistance of the circuit will decrease, and 
total circuit current will increase. Conversely, with 
an open-circuited element, the equivalent resistance 
of the entire circuit will increase while total circuit 
current decreases. 



Wye-Delta Transformations 

Any of the circuits now covered can be reduced to a two- 
terminal network, as seen in figure 2.30A. The circuit 
receives power from an external source and can contain 
resistance, inductance, and capacitance. Such networks are 
called passive. For dc, only resistance is of interest, and it 
can be found from the terminal voltage and current by 



R 



V 
I' 



Numerous circuits can be represented by a two-terminal 
arrangement. Other circuits, including several in mine 
power systems, cannot be represented in this way, but many 
of these can be resolved into the three-terminal network 
given in figure 2.30B Even though with three terminals 
there now exist three voltages and three currents, the con- 
cept of circuit equivalence still holds; that is, voltages and 
currents are identical and the circuits are equivalent. 



Ii 



120V 



a 7H 

t — mh 



6 



isn 



1 */VW- 

lon 




ion 



Figure 2.29. - Series-parallel circuit for example 2.8. 



'12 



I 




1 i— 


__1 


1 « o 




^ 




Figure 2.30.— Two-terminal (A) and three-terminal (B) net- 
works. 



34 



For three-terminal networks, there are two basic circuit 
configurations: the wye (y) and the delta (A) (fig. 2.31). The 
wye is sometimes called a star, but the term y is standard. 
It is sometimes advantageous to replace or substitute the 
three wye-connected resistances with another set that is 
delta-connected, or vice versa. 

By using equivalence of input currents and voltages for 
wye and delta circuits, delta-wye (or delta-to-wye) and wye- 
delta transformations can be derived. Thus for equivalence 
of the circuits in figure 2.31, 



and 



R.» = 



R»c = 



R„ = 



R„R(, + R(,R C "+■ RcR 





Re 




R<.R(> 


+ R R C 


+ R C R„ 




R 


i 


R Rt 


+ R 6 R C 


+ R C R„ 



R* 



(2.45) 



(2.46) 



(2.47) 




Figure 2.31. -Wye (A) and delta (8) circuit configurations. 



In other words, the delta is equivalent to the wye if the 
resistances of the delta are related to the wye by equations 
2.45, 2.46, and 2.47. Accordingly, with three terminals a, 
b, c, containing wye-connected R„, R„, R c , the circuit perform- 
ance is unaffected by replacing them with a delta-connected 
R a6 , R bc , R ca . Likewise, for equivalence of delta-to-wye sets, 



and 



R„ 



R t = 



Rc = 



RotRtc 



R afc 


+ R bc + Rj 




RoiRtc 


R o6 


+ R ic + R c „' 




RacRic 



R„t + Rl,,. + R„ 



(2.48) 



(2.49) 



(2.50) 




© 



R ab 
-WV\A- 



R 



ac 



Ri 



be 



These transformations are useful in allowing three-terminal 
circuit reduction because they allow substitution when a 
network does not contain either series or parallel elements. 
The circuit or circuit portion may not outwardly appear as 
three-terminal, and common examples, n and T, are given 
in figure 2.32. These are actually delta and wye circuits 
drawn in a slightly different fashion. It will be shown in 
chapter 4 that delta and wye circuits are the two most im- 
portant configurations for power systems. These trans- 
formations will be called upon again at that point. 

It has already been shown that when circuit elements 
are neither all in series nor all in parallel, but have some 
other series-and-parallel arrangement, the elements can be 
handled in groups to reduce the circuit to an equivalent 
resistance. This important kind of circuit analysis has been 
called circuit reduction. Now that delta-wye and wye-delta 
transformations have been introduced, the substitution 
process can be employed to solve networks that contain 
elements neither in series nor parallel. A prime instance 
is the common bridge circuit shown in figure 2.33. The 
bridge is one of the most used configurations in electrical 
instrumentation. The objective here is to find all available 
currents and voltage drops in the network, and an overall 
solution approach is illustrated in the following example. 



Figure 2.32. — "T" and V circuit configurations. 



■ab 



ao 



V, 



ab 




bo VWV 



Figure 2.33. — Common bridge circuit. 



35 



EXAMPLE 2.9 

Consider that the resistances shown in figure 2.33 

are as follows: 

R x = 5 Q R 2 = 10 Q R 3 = 15 Q 
R 4 = 20 Q R 5 = 25 Q R 6 = 0.4 Q 

Find the equivalent resistance of the circuit between 

points a and b. 

SOLUTION. The original circuit has been redrawn 
in figure 2.34A, in which a delta configuration is 
clearly defined by points a, c, d. The first step for cir- 
cuit reduction is to convert the delta to a wye. From 
equations 2.48, 2.49, and 2.50, 



R„ 



RadRc 



R ad + R C(J + R 
(10) (5) 



10 + 15 + 5 



= 1.67 Q, 



and 



(5H15) = 25 
30 

Q5H10) = 5£2 
30 



This conversion results in the simple series-parallel 
circuit in figure 2.345. Combining the series elements 
and parallel branches in the center of the circuit fur- 
ther reduces the circuit to that shown in figure 2.34C: 

R c + R 4 = 2.5 + 20 = 22.5 Q, 

R a + R 5 = 5 + 25 = 30 Q, 

= (22.5) (30) = 
22.5 + 30 



The equivalent resistance of the total circuit is then 
simply 

R c , = R„ + R x + R 6 

= 1.67 + 12.9 + 0.4 = 15 £2. 

The total circuit current can now be found using 
Ohm's law; for instance, if 

V ai = 30 V, 



then 



^ = 30 = 2A. 
R.. 15 



Finally, Kirchhoff s current and voltage laws and the 
voltage and current distribution formulas can be 
employed to find currents through and the voltage 
drops across each circuit element. For example, if I 4 
is the current through R 4 , then 



I 4 = U 



= 2(- 



Rd + R 5 



R c + R 4 + Rd + R5 



-) 



30 



22.5 + 30 



-)=1.14A. 



It should be noted that the current through R c is also 
I 4 , but R c does not exist in the original circuit of figure 
2.33. Thus, a problem exists in finding the currents 
through R 1( Rj, and R 3 . One solution would be to solve 
for the three potentials among points a, c, d and use 
Ohm's law to find the three currents in the assigned 
delta connection. 





, R 6 

b )b 

A B 

Figure 2.34. — Circuit reduction of bridge circuit. 



Ft 



Rc 



36 



EXAMPLE 2.10 

Consider that the resistances shown in figure 2.33 

are 

R t = 15 Q R 2 = 15 Q R s = 15 Q 
R 4 = 20 Q R 5 = 25 Q R 6 = 10 Q 

Find the equivalent resistance of the circuit between 

points a and b. 

SOLUTION. Three identical resistances form a delta 
configuration in the circuit, or 

R t = R 2 = R) = 15 Q. 

Following the same processes as in example 2.9 for 
figure 2.34B, 



R„ 



(15) (15) 



= 5 2, 



15 + 15 +15 
R„ = 5 Q 
and R t = 5 Q. 

Now, the center resistance of figure 2.34C is 
(5 + 20) (5 + 25) 



R, 



13.64 Q, 



5 + 20 + 5 + 25 

and the equivalent resistance of the circuit is 

R e , = 5 + 13.64 + 10 = 28.6 Q. 

It should be noted that an important situation is es- 
tablished where all resistances in a delta or a wye 
configuration are equal. If R 4 is each resistance in 
the delta and R y is that in the wye, then from equa- 
tions 2.48, 2.49, or 2.50, 



R v = 



Ra Ra 



or 



Rv = 



Ra + Ra + Ra 
Ra 



The majority of delta or wye configurations used in 
power systems consist of identical elements in each 
leg. 



Much of circuit analysis can be handled by circuit re- 
duction, but as circuits become more complex this process 
becomes cumbersome. Nevertheless, circuit reduction 
should always be used when it produces results more eas- 
ily than other methods. There are solution approaches that 
are more systematic, and the next two sections discuss two 
of these. 

Circuit and Loop Equations 

Before more general solution methods can be identified, 
the meanings of some words need to be clarified. A node 
is the position or point in a circuit where two or more 
elements are connected. When three or more elements ex- 



tend from a node, the node is called a junction. A branch 
is a circuit portion existing between two junctions and may 
contain one element or several in a series. A loop is a sin- 
gle closed path for current. Figure 2.35 illustrates all these 
circuit parts. 

The following technique, loop analysis, is based entirely 
on Ohm's law and Kirchhoff s voltage law. The analysis 
principle produces n simultaneous equations requiring the 
solution of n unknowns, and the unknowns are currents. 

In loop analysis it is only necessary to determine as 
many different currents as there are independent loops; that 
is, the equations are constructed by defining independent- 
loop currents. For example, in figure 2.36, the current 1^ 
flowing out of source V t and through R„, will be around 
loop 1. The current flowing from source V 2 through R c will 
be around loop 2. Although not essential, these directions 
follow the general convention of assigning all reference 
loops clockwise. It is sometimes more desirable to use other 
directions, for instance with currents flowing out of a source 
positive terminal, but it is imperative that the use of cur- 
rents within a specific loop remains consistent after the loop 
is assigned. 

Notice in figure 2.36 that both I t and I 2 flow through 
R fc . Depending on the loop direction, that is, the direction 
defined by I t or I 2 , the total current through R A is either 
I t — 1 2 or I 2 — Ii- Thus if L_ and I 2 can be found, the current 
through each circuit element can be determined. 

The first task in loop analysis is to use Kirchhoff s volt- 
age law to write equations about each current loop, stating 
that the sum of voltages about each loop equals zero. For 
loop 1, 







-V x + RJ, + RA-U 



(2.51) 



Notice that RJ, equals the voltage drop across R„, and 
Rftdj-Lj) equals that across R b . In the latter case, the voltage 
can be taken as RJi — RJ 2 , considering that the voltage 
produced by I 2 opposes that produced by I x . Likewise, for 
loop 2, 



= -V 2 + R C I 2 + R^-U 



Junction - 



Branch 



*-n^-v*-n^f-n^ 



V— *-Hh 



Source- 



^-Nodes /^X 

,r ±: /Loop) <=} Branch J=- 



'7 



(2.52) 



Passive elements 




Figure 2.35. — Parts of circuit. 




Figure 2.36. — Circuit demonstrating two independent loops. 



37 



By rearranging equations 2.51 and 2.52, 



and 



(R„ + RJIi - RJ 2 = V, 
-RJi + (R b + RJI 2 = V 2 , 



(2.53) 
(2.54) 



which are two simultaneous equations with two unknowns, 
Ij and I 2 . These can be solved easily by simultaneous 
methods. 

It should be noted that one additional loop equation 
could be written, that for the loop containing both V! and 
V 2 . However, this will not provide another independent 
equation. Information concerning the maximum number of 
independent equations available will follow shortly. 



EXAMPLE 2.11 

Find the current through the 1.5-Q resistor in 
figure 2.37 using loop analysis. 

SOLUTION. Two loops are defined in the figure 
where the current through the 1.5-Q resistor is 
Ii + I 2 . Applying Kirchhoffs voltage law to loop 1, 

0.51, + 1.5^ + I 2 ) + 1.0I X = 250 
and for loop 2, 

0.5I 2 + 1.5(1! + I 2 ) + 1.0I 2 = 300. 

By simplifying these equations, 

3I X + 1.51a = 250, 
1.5Ii + 3I 2 = 300. 

Simultaneous solution of these results is 

I x = 44.4 A, I 2 = 77.8 A, 

and the current through the 1.5-Q resistor is 

Ii + I 2 = 122.2 A. 



To further enforce the concept of loop analysis, again 
consider the common bridge circuit, which is redrawn in 
figure 2.38 to include current loops. Three loop equations 
can be written because there are three possible indepen- 
dent loops. For loop 1, 

RA - I 2 ) + RA - U + RJ! - V B4 = 0; (2.55) 

for loop 2, 

RA - U + RA +RA - I 3 ) = 0; (2.56) 



for loop 3, 



0.511 



o.sn 

A/WV- 



250 V 




300 V 



Figure 2.37.— Two-loop circuit for example 2.11. 




Figure 2.38. — Bridge circuit demonstrating loop analysis. 

Again, rearranging, 

(R x + R 4 + R^ I x - RJ. 2 - R 4 I 3 = V„ t , (2.58) 

-Rjlj + (R, + R 2 + R 3 ) I 2 - R 3 I 3 = 0; (2.59) 



-R 4 l! - R 3 I 2 + (R 3 + R 4 + R 5 )I 3 = 0; 



(2.60) 



which are three simultaneous equations with three un- 
knowns, I u I 2 , and I 3 . The proper combination of these cur- 
rents will yield the current through each branch of the cir- 
cuit. The process was again to employ Kirchhoffs voltage 
law for the purpose of finding the unknown currents. 

Other loops about the bridge could be assigned and will 
produce the same valid results. Generally, the particular 
choice of loops can enhance a desired result. For instance, 
if only the current through R 3 of figure 2.38 is desired, 
establishing one loop current through that resistor would 
create a more direct solution. 



R 4 (I 3 - Ii) + R 3 (I 3 - I*) + R 5 I 3 = 0. (2.57) 



EXAMPLE 2.12 

Using loop equations, solve for each branch cur- 
rent in the circuit shown in figure 2.39. 

SOLUTION. Applying Kirchhoffs voltage law to 
loops 1 and 2 respectively, 

2(l! - I 2 ) + 5(Ii - I 3 ) + 2I X = 56, 
2(I 2 - I x ) + 10I 2 + 1(I 2 + I 3 ) = 0. 



38 



As the assigned current for loop 3 passes through an 
ideal current source, 

I, = 6 A. 

Therefore, the equations for loops 1 and 2 become 

2^ - I 2 ) + 5(1, + 6) + 21, = 56, 
20, - I,) + 101, + 1(I 2 + 6) = 0, 
or 91, - 2I 2 = 26, 

-21, + 13I 2 = -6. 

Solution of the last two simultaneous equations yields 

I, = 2.9 A, I 2 = A. 

Each branch current can now be resolved from the 
loop currents. For the branch containing the 2-Q 
resistor and the 56-V source, 

I = I, = 2.9 A; 

for the other 2-Q resistor, 

I = I, - I 2 = 2.9 A, 

and the resistors in the other branches, 

J-10O = ±2 = V ■"■» 
1,2 = 1 3 — 1 2 = D A, 

I 4a = I 3 = 6 A, 
I sa = I, + I 3 = 8.9 A. 

A loop equation could have been written for loop 
3, but it can only state that the voltage drops across 
the 1-Q, 5-Q, and 4-Q resistors in that loop are equal 
to the voltage across the 6- A ideal current source, 
which is unknown. Such an equation would only 
complicate the solution to the problem. 



As circuits become more complex and the number of 
possible loops increases, a method for determining the 
number of required equations is useful. By counting the 
number of branches and junctions in the circuit, the follow- 
ing expression provides the necessary number of loop 
currents: 

number of equations = branches — (junctions — 1). 



ion 

^vwv- 



2n C I 2^ in 



o — vwv- 



J V\M " 



56V p Q i 5n (S(t)6A 









2il 4X1 

Figure 2.39. -Three-loop circuit for example 2.12. 



For figure 2.38, there are six branches and four junc- 
tions; therefore, the number of equations needed equals 
6 - (4 - 1) = 3. 

Node Equations 

In the preceding analysis, Kirchhoff s voltage law 
established the method of loop equations. Kirchhoff s cur- 
rent law did not receive any attention, yet it was satisfied. 
This can be demonstrated with figure 2.38 by taking any 
junction and summing the currents through it. Consider- 
ing that the currents through R, and R 2 flow from junction 



or 
hence, 



li l«i 1r 2 — 
I, - (I, - I 2 ) - I 2 = 0; 
I, - I, + I 2 - Ij = 0. 



Kirchhoff s current law is used directly in node analysis, 
and the unknowns are voltages across branches. The tech- 
nique by which these voltages are referenced or measured 
provides a simplifying procedure for a circuit being ana- 
lyzed. Each junction or principal node in a circuit is assigned 
a number or letter. Voltages can then be measured from 
each junction to one specific junction, called the reference 
node. In essence, the reference node is dependent on all 
other nodes in the circuit. Node analysis consists of find- 
ing the voltages from each junction to the reference node. 

The procedure can be demonstrated easily with the 
simple two-junction circuit shown in figure 2.40, in which 
I„, R A , and R c are known. The existing junctions are A and 
O, and O is taken as the reference. The voltage from A to 
O is then Y AO , and Kirchhoff s current law can be used to 
write an equation for junction A: 



L - L - L = 0. 



(2.61a) 



By Ohm's law, 




VylO — IfcRfc 

= IX 


Therefore, 


la 


Rt R c 


since 1/R = G, 


h 


- VaoG„ - V 



= 0; 



G C = 0. 



(2.616) 



(2.62) 



Equation 2.62 can be further solved for unknown, V ao . 

During the process, an equation was written for each 
junction, excluding the reference node. The number of re- 
quired equations for node analysis is therefore always one 
less than the number of junctions in a circuit. 

To illustrate node analysis further, consider the three- 
junction circuit in figure 2.41. If junction is taken as the 
reference node, V AO and V BO are the unknown voltages. The 
reference node, which establishes a reference potential 
across the bottom of the circuit, is normally assumed at zero 
potential. Accordingly, the double-subscripted voltages are 
unnecessary and unknown values can be simply called V^ 
and V B . Further, as zero potential is often referenced to 
earth or ground, a most convenient reference, figure 2.41 
can be redrawn as shown in figure 2.42. These circuit 
elements are still connected to a reference node through the 
ground symbols, as shown. Hence, each of the circuit 
elements is said to be grounded. 



39 



Now, applying KirchhofFs current law to junctions A 
and B, 

L - Ie - l b = 0, (2.63a) 

-h-U+Ic = 0. (2.636) 



By Ohm's law, 
and 



I b = V A G b , h = V B G d 

l = (v„ - V B )G C . 



(2.64) 



The last expression is evident because, by KirchhofFs 
voltage law, the voltage across G= is the potential at junc- 
tion A minus that at junction B. Therefore, 

I. - Wa - V*)G C - V A G b = 0, (2.65) 

-I, - V B G d + <V A - V B )G C = 0, (2.66) 

or (G b + GJV„ - G,V B = I B , (2.67) 

-G e V„ + (G c + G d ) V B = -I e , (2.68) 

which are two simultaneous equations with unknowns, V^ 
and V B . 

The same procedure can be applied to circuits with more 
nodes. The foregoing examples have shown only current 
sources that are known, but node analysis can also be used 
with voltage sources or known voltages. Such is the case 
with figure 2.43, where the current through G a is 



likewise, 



L = (Ya - V B )G a 
Ic = (V e - V B )G C . 



The analysis procedure can then continue as before. Mixed 
voltage and current sources can be handled in much the 
same manner, realizing that the current source establishes 
the current through the branch in which it is contained. 
With both loop and node analysis available, a decision 
must be made as to which technique best suits the solution 
of a circuit. Simply, the one to select is that providing the 
fewest equations to resolve. Since circuit reduction may still 
lead to the most efficient procedure for some circuits, it 
should always be considered. 



EXAMPLE 2.13 

Use node analysis to find the voltage across the 
0.5-Q resistance in figure 2.44. 

SOLUTION. The circuit contains three junctions. If 
the junction at the bottom of the circuit is taken as 
the reference node, A and B can be considered as the 
independent junctions. Here, KirchhofFs current law 
yields 

I A + l AB = 1,500 A, 
I B - lu = 1,000 A. 

The unknown voltages for node analysis are V^ and 
V B , existing between each independent junction and 
the reference node, where 

V, = Ul), 
V B = U2). 

It can also be noted that 

V„ - V B = U0.5). 




Figure 2.40. — Simple two-node circuit. 




Figure 2.41.— Three-junction circuit. 




^- - _ Reference 

Figure 2.42. -Three-junction circuit with grounds. 




Figure 2.43. — Voltage-source circuit demonstrating node 
analysis. 



V 



A ^ B 



I A J L 0.5H II I B 
1.500 A(T) f | in | 2n 



(m- 



OOOA 



Figure 2.44. - Circuit for examples 2.13, 2.15, and 2.16. 



40 



Substituting these Ohm's law relationships into the 
current-law equations produces 

1V„ + 2(V A - V B ) = 1,500, 
0.5V fi - 2(V A -V.) = 1,000. 

Rearranging, 3V A - 2V B = 1,500, 
-2V,4 + 2.5V B = 1,000. 

Solving these two simultaneous equations gives 

V A = 1,644 V, V B = 1,716 V. 

The voltage across the 0.5-Qresistance is then 

Va - V B = -72 V, 

which means that the actual voltage polarity is the 
reverse of that used in the solution and shown in the 
figure. 




EXAMPLE 2.14 

Find the voltage, V lf across the 1-Q resistor in 
figure 2.45 using node analysis. 

SOLUTION. The circuit contains a mixture of cur- 
rent and voltage sources. This presents a difficulty 
for applying node analysis, as the currents associated 
with the voltage sources are not known. However, as 
the objective of node analysis is to find unknown 
voltages, the difficulty can be eliminated by avoiding 
the voltage sources in the solution. This can be done 
by assigning nodes on both sides of each ideal volt- 
age source, treating both nodes and the voltage source 
together, and applying Kirchhoff s current law to both 
nodes simultaneously. For instance in figure 2.45, 
nodes 1 and 2 are on both sides of the 6-V source, and 
nodes 3 and 4 are associated with the 12-V source. 
Each voltage source can be considered a short circuit 
joining its associated nodes, and current flow into the 
combined source and two nodes equals current leav- 
ing the combination. The node-source combinations 
are often termed supernodes and are signified in 
figure 2.45 by the enclosed dashed lines. Each super- 
node reduces the number of nonreference nodes by 
one, thus greatly simplifying the application of node 
analysis. 



6.Q 



12 V i 



Using this concept for the supernode containing 
the 6-V source, Kirchhoffs current law gives 



Ii + Is + le 



2 A. 



Notice that Kirchhoffs current law for the 12-V super- 
nodes produces the same equation. Assigning junc- 
tion 4 as the reference node, the voltages of the cir- 
cuit associated with the nonreference nodes 1, 2, and 
3 are V x , V 2 , and V 3 , where 



Ii 



l 



and 



I =£ 

13 3' 

J „ V 2 ~ Va 



V 2 - V, = V, 
V, = 12 V. 



Rewriting the current-law equation 
1 



V V - V 

2 +1* + Xi la =0 



or 



^6_ 2 + Y 2 + V 2 ^ 
13 6 



12 



= 



or V 2 = 4 V. 

Since V 2 - V, = 6 V, 

the voltage across the 1-S2 resistance is 

V, = 4 - 6 = -2 V, 

which states that the voltage is in the opposite direc- 
tion to that shown in figure 2.45. 



Figure 2.45. - Circuit for example 2.14. 



Network Theorems 

Practically any circuit can be analyzed using either 
circuit reduction, loop equations, or node equations. There 
are also several theorems that allow the simplification of 
particular circuits so that these three methods can be ap- 
plied more easily. The most commonly used theorems are 

• Substitution, 

• Superposition 

• Reciprocity 

• Source transformation, 

• Maximum power transfer, 

• Thevinin's, and 

• Norton's. 

Substitution has already been used extensively and 
simply states that equivalent circuits produce equivalent 
results. The remaining theorems are discussed here. 

Superposition 

The superposition theorem relates that for a linear, bi- 
lateral network with two or more electromotive sources 
(voltage or current), the response in any element of the 



41 



circuit is equal to the sum of responses obtained by each 
source acting separately, with all other sources set equal 
to zero. Although the word "bilateral" is new, it does not 
create problems in dc analysis because passive circuits un- 
der dc are always bilateral. This concept will be discussed 
in more detail later. 

The meaning of superposition can be illustrated using 
figure 2.46A, a network with two voltage sources. The 
theorem relates that 

1. If one source is set equal to zero (removing it from 
the circuit) and the currents produced by the other source 
are found, 

2. Then if the second source is set equal to zero and cur- 
rents caused by the first source are found, 

3. By summing both findings, the results are the cur- 
rents with both sources operating. 

In other words, by letting V 2 = 0, as in figure 2.46J3, through 
circuit reduction, 

R2R3 



R«i — Ri + 
Inn = 



( S 



R 2 + R 

= Vx 
R el 

R 2 



I2U] 



+ R : 
R 3 



R 2 + R 



)Ii[ii> 
)I 1(1 , 



The second part of the double subscripts is used only to 
signify that the currents are caused by source 1. The neg- 
ative sign in the last expression is caused by the current 
direction assumed in the illustration. The next step is let- 
ting V! = 0, thus restoring V 2 (fig. 2.46C), 



R,2 — Rj + 

1:2191 = 
I3U] = (: 
Iim = v 



R1R3 



Ri + R 3 
= V, 

R*2 



Ri + R 
R, 



")l2[2]> 
I 

)I2[2) 



Rj + R 3 
Finally, the sums of steps 1 and 2 yield 

II = IlUl + Il[21> 
*2 = *2[1] + l2[2] > 
I3 = J-311] + J-3I2] » 



(2.69) 
(2.70) 
(2.71) 



which are the currents with both sources in operation as 
in figure 2.34A. The process is adaptable (and perhaps more 
useful) for circuits having more than two voltage or cur- 
rent sources. As with current sources in node analysis, the 
unknowns in each step are voltages. Nevertheless, super- 
position allows many sources to be considered separately, 
and it is of great benefit in the analysis of circuits. 



EXAMPLE 2.15 

Use the superposition theorem to find the voltage 
across the 0.5-Q resistance in figure 2.44. Note that 
this is the same circuit used for example 2.13. 

SOLUTION. Following the first step of the superposi- 
tion theorem, the 1,000- A current source on the right 
side of the circuit will be turned off. The circuit is now 
operating as shown in figure 2.47A. Only L Bm need 
be known to solve the problem. Using current divi- 
sion for the parallel branches, 

l ABll] = 1,500(-^-) = 429 A. 
0.0 

Figure 2.47J3 shows the second step in the problem 
solution, where the 1,500- A source is turned off. Now 
the current through the 0.5-Q resistor is 



L 



1,000(— ) = -571 A. 
3.5 



Summation of these two findings produces the cur- 
rent from A to B with both sources operating. 

•MB = 1>1B[1) T 1/1B[2]> 

l AB = 429 - 571 = 143 A. 



Thus, 



Vab = (-143X0.5) = -72 V. 



It is obvious that this technique produces the answer 
faster than the process given in example 2.13. How- 
ever, node analysis may give a more efficient solu- 
tion with other problems. 



Reciprocity 

The reciprocity theorem states that in a linear passive 
circuit, if a single source in one branch produces a given 
result in a second branch, the identical source in the sec- 
ond branch will produce the same result in the first branch. 

IL Ri r 2 Jl IH!) r i R? fed ! '(2) Rj R2 fe 2 > 

1+ 





Figure 2.46. — Circuit for demonstrating superposition 
theorem. 



1.5QOA A Iab( 1 ) b 



A jABg) B 1 '^ )00A 



^ww- 



05A I 0.5SI 

I- 1,000- A 1,500- A. 



1.500A(t) I A (i)J|in2ni'^e A £urce|in I B(2 )||2n ©'.000A 



off off 



A B 

Figure 2.47. - Circuit In figure 2.44 with sources turned off. 



42 



This reciprocal action is demonstrated in figure 2.48. In 
figure 2.48A, if V, produces I t in the branch that goes 
through R 5 , moving V, to the R 5 branch will produce I 2 in 
the original location of V\ (fig. 2.48B). The currents Ij and 
I 2 will be equal. The dual form of reciprocity has a similar 
function in relating a current source to the voltage pro- 
duced. The great advantage of this theorem is that a source 
may be moved to another location that is more convenient 
to analyze. 

Source Transformation and Maximum 
Power Transfer 

Before defining the theorems associated with source 
transformation and maximum power transfer, it is advis- 
able to expand the topics of ideal and practical sources. An 
ideal voltage source has been defined as a device whose 
terminal voltage is independent of the current that passes 
through it. Although no such device exists in the practical 
world, it is convenient to assume a resistance in series with 
an ideal source as a datum, against which the performance 
of an actual voltage source can be measured. This is shown 
in figure 2.49 where the performance of a 12-V automotive 
storage battery is plotted against an ideal voltage source. 
The internal resistance, R v , compensates the output volt- 
age, V x , for varying load currents, I L . These currents are 
obtained by changing the load, R L . It will be found that with 
small current the practical source approximates the ideal 
one. But under heavy duty where there are high current 
and low load resistance, the output voltage drops substan- 
tially. Using the Ohm and Kirchhoff voltage laws, 



(- 



V, 



Rv + R 



■JRl, 



Il = 



V s 



R v + R/ 



(2.72) 



(2.73) 



V, equals the voltage of the ideal source, which can be found 
by measuring the terminal voltage with no load resistance. 
The internal resistance, R v , can then be determined by 
applying a known R L and measuring V L . 

Similarly, figure 2.50 models a practical current source 
where R, is the internal shunt resistance. The graph illus- 
trates the effect of this resistance: as the load resistance 
increases, terminal current decreases. Using Kirchhoff s 
current law, it can be shown that 



V £ 



R, + R L 



and 



Ir = C 



Rz 



R, + R L 



)I. 



(2.74) 



(2.75) 



The output of the ideal current source, L, can be found by 
short-circuiting the output terminals and measuring the 
resulting current. Then R, can be calculated by measuring 
Vi and l L with a known load, R L . Actually, shorting the ter- 
minals of a source is usually unwise because it could dam- 
age the real-world source, not to mention being an unsafe 
practice. 

I, can also be determined through source transforma- 
tion, which uses the fact that two sources are equivalent 
if each produces identical terminal voltage and current in 
any load. Therefore, for equivalence of practical voltage and 



Ri R 3 

i— M/V- 




A B 

Figure 2.48. - Demonstration of reciprocity theorem. 



O.Olil 




> 12 


1 l ' 
Ideal 


i ' i ' i 
source 










— ilO 


- 




- 


fin 




Practical 




UJ 




source 




£ 6 






- 


h 4 






- 


o „ 








> 2 


/ , i , 


i,i,i 





0.02 0.04 0.06 0.08 0.1 
LOAD (R L ), XI 

Figure 2.49. - Practical voltage-source model. 




t Is 

UJ 

S3 

§G5I S |- 

9 



I 
Ideal 


source 




- 


'- \^ 


Practical 
^ source 


: , 


i i 



R, 2R; 3Rj 

L0AD(R L ) 

Figure 2.50. — Practical current-source model. 



practical current sources, equations 2.72 and 2.73 must 
equal 2.74 and 2.75, respectively. It is obvious that both 
sets are interrelated. In other words, for load current, 



Ix 



R.-L 



R, + Rz R, + Ri 



(2.76) 



If equation 2.76 is valid for any load, R i( it must hold that 

R, = R„ (2.77) 

V. = R,I„ (2.78) 

where R s = the internal resistance for either 
equivalent practical source, 
V 5 = output voltage of ideal voltage source, 
and I, = output current of ideal current source. 

This relationship is shown in figure 2.51. The two circuits 
shown will be named shortly. Source transformation states 
that if one source is known, it can be replaced with the 
other. Note however that even if two practical sources are 
equivalent, the power that the two internal ideal sources 
supply and the internal resistances absorb may be quite 



■■ 



^^■^M 



43 



different. Notwithstanding, this substitution is helpful in 
writing network equations because constant-current sources 
are more convenient for node equations, and constant- 
voltage sources are best for loop equations. In addition, 
the exchange of particular sources may permit direct cir- 
cuit reduction. 



EXAMPLE 2.16 

Solve the problem in example 2.13 using only 
source transformation. 

SOLUTION. Two practical current sources exist in 
figure 2.44 between junctions A and O and between 
junctions B and O. Applying equation 2.78 for the left- 
hand source, 

L, = 1,500 A , 

R rf = 12, 

V*. = 1,500(1) = 1,500 V; 

and for the right-hand source, 

L, = 1,000 A , 

R^ = 2 S2 , 

V* = (1,000X2) = 2,000 V. 

Ri4 V^ and Rsb V^ describe two practical voltage 
sources that can replace the current sources between 
junctions A and O and junctions B and O, respectively. 
Figure 2.52 shows the results of this transformation, 
where the circuit becomes a simple loop. The current 
from A to B is now 

1ab= 1,500- 2,000 = _ 143A> 
3.5 

and the voltage between is 

V AB = (-143X0.5) = -72 V. 

Source transformation also produced results quicker 
than node analysis, but again, this might not occur 
with other circuit configurations. 

In the above solution, practical current sources 
were replaced by practical voltage sources. By com- 
paring figure 2.44 with figure 2.52, it can be seen that 
points A, B, and O exist in both. Caution should 
always be taken to ensure that a desired node is not 
lost after the transformation. 



Since load resistance can vary from zero to infinity, some 
value of resistance must exist that will receive the maxi- 
mum power available from a particular source. It can be 
proven, using the concepts just presented, that an indepen- 
dent voltage source in series with a resistance, R„ or an 
independent current source in parallel with a resistance, 
R„ delivers maximum power to a load resistance, R L , when 
Rx. = R,. This is called the maximum power transfer 
theorem. 

Theveain's and Norton's Theorems 

These theorems are closely related to source transform- 
ation. They can be illustrated by considering the active net- 
work (one that delivers power) with two output terminals 



shown in figure 2.53. Here, the internal configuration is 
unimportant, but the elements must be linear. The sources 
can be either ideal voltage or ideal current. 

Thevenin's theorem states that if an active network 
(fig.2.53A) is attached to any external network (fig. 2.53B), 
it will behave as if it were simply a single ideal voltage 
source, V„, in series with a single resistance, R„ (fig. 2.530. 
In other words, the active circuit will appear as a practical 
voltage source. Values for V„ and R„ can be found as follows. 
When all internal sources are operating normally and no 
loads are connected, the open-circuit voltage across the out- 
put terminals equals V„. With all the ideal sources turned 
off, a resistance, R,,, can be measured at the terminals. This 
is because when an ideal current source is turned off, it ap- 
pears as an open circuit (an infinite resistance). An ideal 
voltage source that is not operating acts as a short circuit, 
thus having zero resistance. 

This theorem is important because it means that any 
linear circuit where the internal components are unknown 
can be considered as a constant-voltage source in series with 
a resistance. Any circuit reduced to this form is called a 
Thevenin circuit. 

Norton's theorem is the corollary to Thevenin's theorem. 
Norton relates that if such an active network is attached 
to any external network, it will behave as a single ideal 
current source, I , in parallel with a single resistance, R„. 

The values for V„ and R„ can be determined by consider- 
ing the same linear active network, this time as showrrin 
figure 2.54A, with internal sources operating normally. The 



-Wh o 

Re 



v s =i s Rs 




Figure 2.51. -Source transformation. 



: AB 

in a — ► b 2n 

-A/WY— *- J WV\r-«— vwv- 

0.5 n i + 

1,500 V O O 2,000V 






Figure 2.52. - Circuit In figure 2.44 with current sources 
transformed to voltage sources. 





+ 


Active 
network 









1 




Active 
network 




External 
network 









A B 

Figure 2.53. -Thevenin's theorem 




44 



output terminals are short-circuited, and a terminal cur- 
rent is measured to give the value for t,. R„ is found in ex- 
actly the same way as in Thevenin's theorem. The combina- 
tion of these elements gives the practical current source 
shown in figure 2.54C, which is also known as a Norton 
circuit. 

The Thevenin and Norton circuits are obviously related 
by source transformation so that if one is known, the other 
can be constructed. The equations relating the two are 
shown in figure 2.55. These theorems are usually employed 
when a series of calculations involves changing one part 
of a network while keeping another part constant. This 
manipulation helps to simplify complex computations such 
as power-system short-circuit currents. 



EXAMPLE 2.17 

Find the Thevenin and Norton equivalents for the 
circuit shown in figure 2.56. 

SOLUTION. Applying Thevenin's theorem, the 
equivalent resistance of the circuit between a and b 
with the internal source off is R„. When the 50-V 
source is off, it acts as a short circuit, shorting out 
the 50-Q resistance in parallel with it. Thus, 



R = R ab = 2 + 



(10) (10) 
10 + 10 



7 Q. 



The voltage across a and b with the internal source 
operating is V„. Using circuit reduction, the equiva- 
lent resistance as seen by the 50-V source with no load 
across the terminals a and b is 



R.,= 



50 (10 + 10) 
50 + 10 + 10 



= 14.3 Q. 



(Note that this resistance is not R„.) The current 
delivered by the source is 



Ix = 



50 
14.3 



= 3.5 A, 



and from current division, 
50 



I 2 = (3.5) — = 2.5 A. 
70 



As no current is flowing between terminals a and b, 
V 2 shown in figure 2.56 is equal to V ai , which is equal 
to V„. Thus, 

V = V a6 = V 2 = (2.5) (10) = 25 V. 

V„ and R„ describe the Thevenin equivalent, and L and 
R represent the Norton equivalent where 

L = ^ = 25 = 3.6A. 
Ro 7 



ALTERNATIVE SOLUTION The definition for R, 
in Norton's theorem is the same as in Thevenin's, 
again, 

R„ = R„6 = 7 Q. 



However, Norton states that if the terminals a and 
b are short-circuited, the current through that short 
circuit is X,. The short circuit is noted by the dashed 
line in figure 2.56. Using circuit reduction, the 2-Q 
resistance connected to terminal a is in parallel with 
the 10-Q resistance connected to terminal b, or 



(10) (2) 
10 + 2 



= 1.67 Q. 



The equivalent resistance as seen by the 50-V source 

is 

(50) (10 + 1.67) = 
50 + 10 + 1.67 

and the current from the source is 



Ix = 



50 



9.46 
From current division, 

I 2 = (5.28) ( 



= 5.28 A. 



50 



50 + 10 + 1.67 



) = 4.28 A, 



and the current through the shorted terminals is 

10 



I. 



(4.28) (- 



-) = 3.6 A. 



10 + 2 
R„ and !«, again describe the Norton equivalent. 



















Active 
network 


X 


K 


Active 
network 




External 
network 


ioCpRoi 


External 
network 

° 


_J 






° 




~^ 







B C 

Figure 2.54. — Norton's theorem. 



VioRo 



I =^ 

l o EL 



0_J 


Active 
network 




i 



= v, 




= Io(D i R o 



Figure 2.55. - Comparison of Thevenin's and Norton's 
circuits. 



2n 

AAM — ?a 



50 V 




Figure 2.56. - Circuit for example 2.17. 



45 



EXAMPLE 2.18 

Determine the Thevenin's and Norton's equiva- 
lents for the circuit in figure 2.57. 

SOLUTION. In the branch containing the 900-V 
source, the two 5-Q resistances are in series. If these 
are combined into one 10-Q resistance, it should be 
quite obvious that two practical voltage sources ex- 
ist between junction 1 and the junction connected to 
terminal b. Source transformation can be employed 
to solve the problem. The resistance and magnitude 
of the ideal current source of the Norton equivalent 
to the 900-V and 10-Q source are 

R t = 10 Q, 

I 1 = ^0 = 90A. 
10 

For the Norton equivalent of the 2,250-V and 15-Q 
source, 

Rj = 15 S, 

j = 2,250 = 15Q A 
15 

Figure 2.58A shows the voltage sources transformed 
to practical current sources. Notice that junction 1 and 
the junction associated with terminal b still exist. Be- 
tween these two terminals, the 90- and 150-A sources 
are operating in parallel, and the 10- and 15-Q 
resistances are connected in parallel. Combining these 
ideal current sources and resistance results in the cir- 
cuit of figure 2.58S. Again, notice that the afore- 
mentioned junctions are retained. Converting the 
60-A and 6-Q current source to its Thevenin equiva- 
lent produces the circuit in figure 2.58C. The 6- and 
4-Q resistances in series with the 360-V are combined 
in figure 2.58D. The 360-V source and 10-Q resistance 
form a practical voltage source between terminals a 
and b, and this is converted to its Norton equivalent 
in figure 2.58.E. Here, simple combination of the two 
parallel 10-Q resistances yields one answer to the 
original problem and is shown in figure 2.58F. The 
remaining answer, the Thevenin equivalent, is in 
figure 2.58G, obtained by source transformation of 
figure 2.58F. 



511 1 411 

-tyW\r f ° a 



To summarize the preceding sections, the fundamental 
laws and parameters were first applied to circuits under 
the influence of dc. Expanding upon these laws, several 
circuit-analysis techniques and theorems were covered. Be- 
cause only dc was considered, resistance was the only cir- 
cuit element of interest. As will be shown shortly, most of 
this theory is also valid for circuits acting under current 
forms other than dc, where inductance and capacitance may 
also enter into the picture. 

TIME-VARYING VOLTAGES AND CURRENTS 

As the name implies, the magnitude of time-varying 
voltages and currents may not be constant with time. Con- 
sequently, the instantaneous values of the voltage and cur- 



900 V | 




Figure 2.57. -Active circuit for example 2.18. 



1 4A 

-WVW 



-OO 



-od 



I)9oa |ion (t)i50A fisn §ion 

^b 



T)36A |io.n. lion 

I 1 1 — ob 



i 4n 

-awv- 



60 A (f 




36A(t) | 5 n 



^>b 



6n 1 4A 



360 V 




180 V( 



sn 

-WA od 



^>b 




360 V 



Figure 2.58. — Circuits illustrating solution steps to example 
2.18. 



rent waveforms, v and i, must be considered. Both v and 
i are functions of time, as they were when originally intro- 
duced in this chapter, and they can assume any form from 
constant to the most complex. Figure 2.59 presents just a 
minor sampling of time-varying waveforms to illustrate 
their general characteristics. 

As with dc, the method for analyzing circuits that have 
time-varying current and voltage is first to form a model 
of the circuit, then to apply the fundamental laws and rela- 
tionships. Unlike dc circuits, a differential equation usu- 
ally results. To demonstrate the effect of time-varying and 
current on circuit elements, this section will first consider 
a special waveform, steady alternating current (ac). 



46 



An example of a steady-state ac waveform is provided 
in figure 2.60. The repetitive nature of this sinusoidal func- 
tion can be expressed mathematically as 



i = I„cos(cut) 



(2.79) 



where 



= current at any time, t, 

= crest or maximum value of current, a 

constant 
= radian frequency, rad/s. 



The term sinusoid or sine wave is used collectively to in- 
clude cosinusoidal or cosine-wave expressions. The above 
equation could also use a sine function, but the cosine is 
employed for convenience when referring to current. 

It can be seen in figure 2.60 and equation 2.79 that the 
instantaneous value of current repeats itself every 2n rad 
or 360°; that is, the waveform goes through one complete 
cycle every 2n rad. The number of cycles per second is cu/2tt 
which is defined as the frequency, f, of the waveform or 



f = 



or 



2n 



CD = 2nf. 



(2.80) 



(2.81) 



The units of frequency are hertz (Hz). One hertz is equal 
to 1 cycle-per-second (cps), an expression whose use is now 
obsolete. The common power frequency in the United States 
is 60Hz, for which go = 377 rad/s, or just simply co = 377. 
A more general form of ac is 



where 



i = I m cos(cut + 6), 
6 = phase angle. 



(2.82) 



Instead of expressing the phase angle in radians, such as 
rt/6, angular degrees, 30°, are customarily used. By adjusting 
9, the sinusoid can be moved left (increasing 0) or right 
(decreasing 0. Such movement is illustrated in figure 2.61. 
Using the earlier technique of developing differential 
equations through circuit analysis, steady ac can be applied 
to pure resistance, inductance, and capacitance to observe 
what happens. 

Alternating Current Through Resistance 

Figure 2.62A shows a resistor of resistance R. From 
equation 2.79, if the current through this element is 

i = I m cos(cut), 

by Ohm's law, the voltage developed across the resistor is 

v = Ri 



or 



v = R(I m cos(cut)) = V m cos(o)t), 



(2.83) 



where V m = RI„ 



= maximum or crest value 
of voltage waveform, V. 



Figure 2.62B shows both voltage and current as functions 
of time. At every instant, v is proportional to i, and v and 
i are said to be in phase. When two sinusoidal waves are 
compared for phase in this manner, both must be sine waves 
or cosine waves; both must be expressed with positive amp- 
litude and have the same frequency. 







f 7T 27T 



2 u 2 



| ; 2 v 

2 



-cut 



V 







■cut 



7T 2tt 37T 

Figure 2.59. - Some time-varying electrical waves. 





-^m 


a 


/ !b 


\ 
\ 


/ 
; 


\ 


T/2 / 


T 


t in seconds 


; 




i W 1 


277" 


cut in electrical radians 






\ 180° / 


360° 


cut in electrical degrees 



Figure 2.60. - Sinusoidal ac waveform. 



i = I m cos cut 




Figure 2.61. -Steady ac showing phase shift. 



+ A 




A B 

Figure 2.62. — Steady ac through resistance. 



47 



Alternating Current Through Inductance 

Suppose that current through the pure inductance of 
figure 2.63A is again as in equation 2.79. The voltage across 
the element is 

T di 

v = L — 

dt 



or 



or 



v = L — (I m cos(o)t)) 
dt 



v = LI, 



dcos(a)t) 

' dt 



Differentiating, v = -a>LI m sin(cot) 

or v = wLI m cos(cot + 90°) = V m cos(ot + 90°), (2.84) 

where V m = coLI m = maximum or crest voltage. 

The term o>L is used so frequently that it is provided with 
a special name, inductive reactance, and is designated "X," 
where 



and 



X = coL =2nfL 
V m = L.X. 



(2.85) 
(2.86) 



Figure 2.63B compares equations 2.79 and 2.84, with i and 
v as functions of time. Here, it can be seen that the current 
crest is reached at a later time than the crest voltage. The 
current waveform is said to lag the voltage waveform by 
90°. The phase angle is called lagging. 

Alternating Current Through Capacitance 

Consider the capacitance shown in figure 2.64A, and let 
the voltage across it be 



v = V m cos(a>t). 



(2.87) 



The current through the capacitor is then 

,dv 

'dt 



i = C^ 



or 



i = C-^-(V m cos(a)t)). 
dt 



Differentiating, i = -a>CV„,sin(o)t) 

or i = o)CV m cos(cot + 90°) = I m cos(cot + 90°), (2.88) 

where I m = coCV m = maximum or crest current 
through the capacitor. 

As with the inductive resistance, coC is also provided a 
special name, capacitive susceptance, and symbol, "B." Thus, 



and 



B = coC = 2TtfC 
L= BV m . 



(2.89) 
(2.90) 



The relationship between the current and voltage wave- 
forms (fig. 2.64B) is the reverse of the inductance situation; 
the current waveform is now leading the voltage waveform 




A B 

Figure 2.63. -Steady ac through inductance. 




Figure 2.64. — Steady ac through capacitance. 



by 90°. The phase angle is also called leading. The impor- 
tance of current and voltage waveforms being compared for 
lagging and leading phase angles will be brought out later 
in this and the next two chapters. 

Time-Varying Equations 

The preceding discussion considered voltage and cur- 
rent to be steady sinusoids. But what if they are allowed 
to have any form? To illustrate the consequences, the fun- 
damental laws and parameters can be applied to the sim- 
ple series RL, RC, and RLC circuits shown in figures 2.65, 
2.66, and 2.67, respectively. 

For the series RL circuit, using Kirchhoff s voltage law 

V = V* + Vt. 

Substituting in the relationships for voltages across resist- 
ance and inductance, 



v = iR + L 



'dt' 



(2.91) 



Now for the series RC circuits, 

v = v R + v c . 

Applying the elementary laws, 

v = iR + — /' idt + V . 
C o 



(2.92) 



The differential equations 2.91 and 2.92 are valid for any 
voltage and current, no matter what form. As before, V 
is the initial charge on the capacitance. 

Considering figure 2.67, which shows the series RLC 
combination, 



thus, 



v = v* + Vi + v c ; 

v = iR + L^- + - /' idt + V . 
dt C » 



(2.93) 



48 



To arrive at an equation that is easier to handle mathe- 
matically, equation 2.93 can be differentiated once: 



dv = R di_ + L d^ + 1 Q 
dt dt dt 2 C 



(2.94) 



This equation again describes or models the circuit for all 
electrical situations, as no restrictions have been placed on 
voltage and current. 

The preceding has shown that when voltages and cur- 
rents represent any form, the application of circuit rela- 
tionships results in a differential equation. Through clas- 
sical differential-equation methods, such equations can 
provide the required solution, but these techniques will not 
be shown because it can confuse the understanding of the 
vital aspects of electrical fundamental methods. 



6 




Figure 2.65. — Simple series RL circuit. 



V Q 




Transients and Circuit Response 

Solution of these equations for all situations yields the 
complete response of the circuit. For linear circuits, the solu- 
tion will have two parts: forced response and natural 
response. The forced or steady-state response can be attrib- 
uted directly to the applied source or forcing function. This 
is the action of voltage and current within the circuit if no 
changes or disturbances are made. The natural or transi- 
ent response is a characteristic of the circuit only, not a 
result of the sources. Such action occurs when a circuit is 
disturbed by a change in the applied sources or in one of 
the circuit elements. After the change, the circuit currents 
and voltages undergo transition from their original state 
to the point where their action is again steady state. The 
time period involved is normally very short, and the occur- 
rence within the transition is called a transient. 

For simplicity, the forcing functions mentioned earlier 
in this chapter were dc, and in network analysis the study 
was devoted only to resistive circuits and dc sources because 
here only the forced response is present. When both induc- 
tance and capacitance are circuit elements, both forced and 
transient responses can be encountered. However, knowl- 
edge of circuit transients is not required when considering 
steady-state voltages and currents, as was seen in the case 
of steady ac. By far the majority of mine power problems 
only require knowledge of steady-state circuit currents and 
voltages, and it will be shown that even though inductance 
and capacitance might be present, as long as only the 
steady-state response is considered the solution of differen- 
tial equations is not needed. However, transient circuit 
responses are an extremely important input in the design 
of mine power systems, and they will be explained in de- 
tail in_chapter 11. 

It has been shown in this section that any resistor, in- 
ductor, or capacitor carrying a sinusoidal current has a 
sinusoidal voltage developed across it. Furthermore, the 
sum or difference of two sinusoidal waveforms with the 
same frequency is another sinusoid. From these concepts, 
it can be shown that for a steady-state circuit, if voltage 
or current at any part of a linear circuit is sinusoidal 
(alternating at a particular frequency), voltages and cur- 
rents in every part of the circuit are sinusoidal with the 
same frequency. 



Figure 2.66. — Simple series RC circuit. 



STEADY ALTERNATING CURRENT 




The form of steady ac has already been shown and used 
in the analysis of simple ac circuits, but here the concepts 
of steady-state ac circuit analysis are introduced. This 
necessitates a review of a familiar but easily forgotten sub- 
ject, complex algebra. 

Real numbers such as 2, 4, and n are easy to understand 
in terms of physical things. Any mathematical operation 
on these numbers always results in another real number, 
except when the square root of a negative real number is 
taken. The term \/ -1 cannot be satisfied by any real 
number. Therefore, the square root of any negative num- 
ber is called an imaginary number. Mathematicians dis- 
tinguish imaginary numbers by writing "i" in front of them, 
but to avoid confusion with the symbol for current, electri- 
cal engineers use the symbol "j" where 



Figure 2.67. -Simple series RLC circuit. 



49 



Addition or subtraction of imaginary numbers yields 
another imaginary number. Yet, when an imaginary num- 
ber is added to a real number, a complex number is created. 
These have the rectangular form, x + jy (for instance, 
3 + j4), where x is the real part and y the imaginary part 
or if 



then 



Z = x + jy, 
Re[Z] = x Im[Z] = y. 



(2.95) 



Complex numbers can be represented graphically by a 
pair of perpendicular axes as shown in figure 2.68. The 
horizontal axis is for real quantities, the vertical one for 
imaginary. Considering x + jy, if y = 0, the complex num- 
ber is a pure real number and falls somewhere on the real 
axis. Similarly, if x = 0, the complex number (now be- 
ing purely imaginary) exists on the vertical axis. Hence, 
complex numbers encompass all real and all imaginary 
numbers. 

In the case of the rectangular forms 

Z = x + jy, 
W = u + jy, 

the following common definitions and mathematical opera- 
tions of complex algebra are applied. 

1. Two complex numbers are equal if and only if the 
real components are equal and the imaginary components 
are equal: 

Z = W, IFF x = u, y = v. 

2. To sum two complex numbers, the real and imaginary 
parts are summed separately: 

Z ± W = (x ± u) + j(y ± v). 

3. The product of a real and an imaginary number is 
imaginary: 

x(jy) = j(xy). 

4. The product of two imaginary numbers is a negative 
real number: 

(jy) (jv) = -yv. 

5. The multiplication of two complex numbers follows the 
rules of algebra (note, an easier way to perform the multi- 
plication will be shown): 

(x+jy) (u+jv) = xu -I- jxv + juy - yv 

= (xu - yv) + j(xv + uy). 

6. By definition, the conjugate of a complex number is 
formed by changing the sign of the imaginary part. An 
asterisk denotes the conjugate: 



Z = x + jy 



becomes 



x - jy. 



7. For division, the numerator and denominator are 
multiplied by the conjugate of the denominator (again, an 
easier method exists): 

x +jy = x + jy (u + jv) = ( xu - yv j + j( uy + xv } 
u — jv u — jv (u + jv) (u 2 + V 2 U 2 + V 2 

Besides the rectangular, there are three other general 
forms of complex numbers: trigonometric, polar, and ex- 



ponential. Figure 2.69 illustrates the conversion of rectan- 
gular to trigonometric or polar forms where 

Z = x + jy. (2.96) 

The absolute value of Z is represented by "r," and 

x = rcos0, 
y = rsin0, 



= tan" 1 (£), 
x 

r = (x 2 + y 2 ) 1/2 . 



where 

x 

(2.97) 
Hence, the trigonometrical form of the complex number is 

Z = r(cos0 + jsin0), (2.98a) 

with the conjugate 

Z* = r(cos0 - jsin0). (2.986) 

The polar form is widely used in circuit analysis and is sim- 
ply written as 



Z = r|0 
and the conjugate, 

Z* = r|H? 
Euler's theorem states that 

cos0 + jsin0 = e' 9 . 



(2.99a) 



(2.996) 



(2.100) 







J3 


— 












J2 


- 








1 


I 


J1 

i , 




I 


1 


1 


-3 


-2 


-1 
-J1 
-J2 
-i3 




1 


2 


3 



Figure 2.68. -Graphical representation of complex number. 




REAL (y) 



Figure 2.69. —Trigonometric or polar representation of com- 
plex number. 



50 



This expression allows a complex number to be written as 
an exponent, the exponential form, 

Z = r(cos0 + jsin0) = re> 8 (2.100a) 

and Z* = re ■*. (2.1006) 

All four complex forms are therefore identical or 

Z = x + jy = Kcos0 + jsin0) = r|0 = re' 8 . 

The form should be selected that gives the easiest 
mathematical manipulation of complex numbers. For ad- 
dition or subtraction, the rectangular expression is best, but 
multiplication and division are much more convenient when 
the number is in exponential or polar form, the latter be- 
ing the most used. For instance, in polar, 

Z,Z 2 = rjfl r a |+ = r x r g | fl + <t> , 

Z 2 r 2 |j> r 2 
and in exponential, 



ZiZ 2 


= 


(r,e^)(r 


,e*) = 


= r 1 r 2 e /(8 


-♦) 




z, 


r^ 8 
r 2 e" 




£,(»-.)_ 





It will be shown shortly that circuits containing resist- 
ance, inductance, and capacitance can be represented by 
complex numbers, and that the solution of these circuits 
under steady ac will use complex algebra. This can be done 
with almost as much ease as the dc circuit analysis pre- 
sented earlier. 



EXAMPLE 2.19 

Find the answer to the following expression in 
polar and rectangular form: 

(2 + j6)(18 |21°) . 
(1.63JM2.6 + jl) 

SOLUTION. Both the numerator and denominator 
of the above expression are multiplications of complex 
quantities. For ease of solution, the rectangular term 
should be converted to polar. This results in 

(6.32|71.6°)(18|21°) 

(1.63|90°)(2.79|21 o ) 

(6.32) (18) 



or 
or 



(1.63) (2.79) 



71.6° + 21' 



90° - 21° 



251-18.4°. 



Effective Alternating Current 

The power available in the outlets of U.S. homes is a 
very familiar quantity: it is sinusoidal, having a frequency 
of 60 Hz and a voltage of 115 V. But what does 115 V actu- 
ally stand for? 



The voltage waveform, being a sinusoid, is not constant 
with time. Therefore, the voltage is certainly not instan- 
taneous. If a measuring device could be connected to an 
outlet in order to visually observe the waveform, it would 
be found that "voltage" is not the maximum value, V„„ be- 
cause this waveform crest is HSyT or 162.6 V. "Voltage" 
does not describe an average value either, because the 
average of a sine wave is identically zero. As another resort, 
the average throughout one positive or one negative half- 
cycle of the waveform could be calculated, but the result 
gives a measurement of 0.637 V m or 103.5 V. To discover 
the meaning of the term voltage, the reason for measuring 
the voltage must be considered. In any system, current and 
voltage are defined in terms of what they will do. Conse- 
quently, the voltage is the effective value of the sinusoidal 
waveform. It is a measure of the effectiveness of the volt- 
age source in delivering power to a resistive load. The 
effective value is called root-mean-square (rms). 

In order to understand rms measurements, it is neces- 
sary to return to the concept of instantaneous power, where 



If the power was being developed across a resistance, R, it 
was shown that 



and 



p = i 2 R 

v 2 
P= R- 



These equations have little practical value for ac as they 
represent the value of power for a particular instant and 
in ac this is ever changing. A more effective measure for 
the value of power is based on the fact that power is the 
rate of doing work. A reasonable measure is then the ave- 
rage rate or average power. For average power, P, consumed 
by the resistance, R, 



P = ave(p) = ave(i 2 R) = (ave i 2 )R 



and 



Pv 
= ave — 
R 



(ave v 2 ) 



Average power is then an effective way to measure or 
quantify ac voltage and current. It has already been seen 
that the units of voltage and current in dc are easy to com- 
prehend; the magnitudes are constant with time, and their 
ability to deliver power is constant. Therefore, it is appro- 
priate to equate ac and dc rates of work, P oc and P<, c , re- 
spectively, in order to determine an effective measurement 
for alternating voltages and currents: 

p dc =PR = P oc = (ave i 2 )R 



or 



I 2 = (ave i 2 ) 



or I = V (ave i 2 ) = ™s current. 

Employing the same procedure, 



V = \/(ave v 2 ) = rms voltage. 



(2.101a) 



(2.1016) 



Current and voltage in ac are therefore expressed as the 
square root of the mean-square values or rms. They are 
sometimes written I rmj and V rTO . It can be shown from the 
voltage and current waveforms (that is, substituting 



51 



I m cos(cut + 9) into equation 2.101a and similarly for voltage) 
that 

I™ = —= or L = \T~2~ !"-. (2.102a) 



and 



V rms = 



v„ 



or V. 



(2.1026) 



Root-mean-square currents and voltages are used so 
often that they are directly implied when referring to an 
ac magnitude. They are almost always used in calcula- 
tions. For simplicity, the subscripts of V rmj and l rms are 
eliminated in practice, and just V and I are written to in- 
dicate rms voltages and currents. All common ac voltmeters 
and ammeters are also calibrated to read rms values. 

The preceding analysis of average power concepts ap- 
plies only to resistance. Average power in the steady state 
supplied to either a theoretically pure inductance or pure 
capacitance is identically zero. This can be proved by inte- 
grating instantaneous power to these elements to obtain 
an average. The results show that the energy received dur- 
ing one-half cycle is stored and then transferred back to the 
source through the balance of the cycle. The stored energy 
in the capacitance is greatest at the maximum of the volt- 
age wave, while in the inductance it is maximum at the 
current-wave crest. 

Phasors 

A steady-state sinusoidal current or voltage at a given 
frequency is characterized by only two parameters: ampli- 
tude and phase angle. This can be seen in figure 2.70A, 
which shows two voltage waveforms separated by a phase 
angle. An ac quantity may also be represented graphically 
by a phasor, illustrated in figure 2.70B. The phasor is a 
continually rotating line that shows magnitude and direc- 
tion (time). In this figure, the phasor is assumed to have 
a length representative of V m , rotation about point 0, and 
an angle increasing with time according to 9, = cot + 9. The 
figure shows the line as if a snapshot had been taken, freez- 
ing action. The alternating quantity, V m cos(a>t + 0), is the 
projection of the phasor on the horizontal axis. In other 
words, as the phasor in figure 2.70B rotates, a plot of its 
projection on the horizontal axis with time reproduces the 
waveform in figure 2.70A. The phasor length shown here 
represents crest voltage but does not necessarily need to 
be equal to it. It is common practice to draw phasors in 
terms of effective (rms) values. Although voltage has been 
employed as an example, phasors can also represent sinu- 
soidal current, among other things. 

Voltage and current phasors are both illustrated as 
rotating lines in figure 2.7 1A, where 

v = V m cos (cot + <{>), 



For example, in figure 2. 7 IB the phasor is shown where 
the current phasor angle is zero. Here, the current phasor 
is termed a reference phasor, and all other phasors are 
drawn relative to it. Either voltage or current can be 
selected as the reference. 

A phasor may be expressed in several ways. To illus- 
trate the most used expressions, consider figure 2.72A, 
which shows one phasor displaced from the horizontal by 
an angle, cot + 9. Recalling complex algebra, the horizon- 
tal axis can be assigned as a real_axis and the vertical as 
the imaginary axis. The phasor, V, is then the sum of the 
real and imaginary components, 



or 



V re and V lm , 
V = V„ + V im . 



(2.103) 




V m cos(cut+0) 

\ \ J I \ \ I I K 

Phase 
angle 

A B 

Figure 2.70.— Sinusoid versus time (A) and as phasor (B). 



<u>\+8 





B 



Figure 2.71.— Phasor representation of current {A) and 
voltage (B). 



i = I„cos cot. 

To show both current and voltage, two phasors can be 
drawn, with one of them advanced by the phase angle, <{>. 
Both lines rotate indefinitely about the axes, and one line 
will always lead the other in the same relative position; 
therefore, the axes are superfluous and need not be drawn. 
Since it is necessary to orient the phasors at a specific 
point in time, a convenient instant is selected as a reference. 




Vj m =Vsin(a>t + 6) 




A B 

Figure 2.72.— Other expressions for phasors. 



52 



Figure 2.12B clearly illustrates the rectilinear form of 
equation 2.103. The real and imaginary components of the 
phasor are 

V„ = Vcos(cot + 0), (2.104a) 

V )m = jVsinM + 0), (2.1046) 

Thus V = Vcos(a)t + 0) + jVsin (cot + 0) (2.104c) 

or V = V(cos(a>t + 0) + jsin (cot + 6)). (2.104<f) 

V is used to signify that the voltage is a phasor ,_and once 
more the imaginary operator, j, signifies that V im exists 
on the imaginary axis. Accordingly, the phasor may be 
considered as the vector sum of two phasors at right angles 
to each other. Applying Euler's theorem (equation 2.100) 
to equation 2.104d. 



or 



V = Ve""" + "> 
V = Ve^'e". 



(2.105a) 
(2.1056) 



The factor, e""', is superfluous, as it contains no unique 
information about the phasor, and it can be suppressed: 



Ve". 



(2.106) 



This is called the exponential form of the phasor. Thus equa- 
tion 2.105c can be expressed in polar form, 



V 10. 



(2.107) 



These phasor forms are very useful in solving ac cir- 
cuit problems. The terms phasor and vector are often 
interchanged. 



Phasors and Complex Quantities 

When introducing the action of time-varying sinusoids, 
certain voltage-current phase-angle relationships were 
found to exist for pure resistive, inductive, and capacitive 
circuit elements. In general, if a steady-state sinusoidal 
current has the time-domain form 



and voltage, 



I = I m cos(cot + 0), 



v = V m cos(cot -I- <}>), 



(2.108a) 



(2.1086) 



current is said to be lagging voltage by the phase angle, 
$— (or conversely, leading voltage by the phase angle, 
0—$). Using the exponential and polar phasors, this current 
and voltage can also be stated 

I = Ie""' + •>, I = Ie", or I = I|0 (2.109a) 

and V = Ve""' + •», V = Ve", or V = I|$ (2.1096) 

where I and V = rms values of current and voltage, 
respectively. 

Before steady-state circuit analysis can be performed, 
pure circuit elements must again be considered, this time 
to analyze the voltage-current relationships using complex 



quantities. Equations 2.108a and 2.1086 are assumed to 
represent the general current through and voltage across 
each element. 



EXAMPLE 2.20 

A circuit has the following voltage and current 
waveforms applied across and through its terminals: 

v = 282.8 cos (377t - 20°), 
i = 42.4 cos (377t + 25°). 

Write the phasor expression for voltage and current. 
What is the phase angle between current and voltage? 

SOLUTION. The two given expressions are in the 
time domain, where for the voltage, 

V„ = 282.8 V, 
4> = -20°, 

and for the current, 

I m = 42.4 A, 
= 25°. 

The phasors for voltage and current are then, 
respectively, 



v v - 



I = 






i = 200 1-20° V, 



/T 



= 30 125° A. 



The current waveform is leading the voltage wave- 
form, and the phase angle between current and 
voltage is 

4> - = -20° - 25° = -45°. 



If sinusoid current is applied to a resistance, R, the 
voltage across it is 

v = Ri. 

Applying the general time-domain expressions, 

V m cos(cot + 4>) = RI m cos(cot + 0) 
or in exponential form, 

y e ir(»rt») _ RIe><"" + <». 

Suppressing e""', 

Ve* = Rle", 
or in polar form, 

V 14 = RI |0. 



53 



In phasor form, V |£ and I |0 are the phasor polar 
representations, 



V = RI. 



(2.110) 



This is the same relationship that exists for time-varying 
waveforms and dc. It is apparent that angles and $ are 
equal and that voltage and current are in phase (fig. 2.73A). 
Suppose the same general forms of current and voltages 
were applied to a pure inductance where, as before, 

dt 

then, using the general exponentials, 

Ve"»<+*> = L — (Ie""" + '"). 
dt 

Differentiating (e"' is a constant with time), 

Ve/ (■«♦•> _ jajLIe"""* 9 ', 

and suppressing &"", 

Ve* = juLIe". 

Thus, in phasor form, 

V = ja>LI. (2.111) 

The imaginary operator, j, denotes a +90° displacement of 
voltage from current; such as illustrated in figure 2.73J3. 
In general, if the current phasor has an angle, 0, the voltage 
phasor angle, <(>, is + 90° for a pure inductance. 
For a pure capacitance, 

dv 



i = C 

dt 

Employing the same process to find equation 2.111, 

I = jwC V. (2.112a) 



or 



v = (t^-)! 



(-£D. 



(2.1126) 



In this case, -j indicates a —90° displacement of the voltage 
phasor from current, as shown in figure 2.73C. 

Now that the phasor relationships of the fundamental 
elements have been covered, the stage is set for impedance 
transforms. 



A very important quantity, impedance, signified by Z, is 
defined as the ratio of the phasor voltage to the phasor cur- 
rent for a circuit or 



Z = 



(2.113) 



This expression is often called Ohm's law for ac circuits. 
Impedance is a complex quantity with dimensions of ohms, 
but it is not a phasor. Therefore, the impedance of the pure 
passive circuit elements, resistance, inductance, and 
capacitance, are respectively 



Zr — R, 



jcoL, Z c = 



jcoC 



(2.114) 



These can be applied directly to circuit analysis when a cir- 
cuit is in steady state. In other words, element impedances 
are employed to convert or transform a time-domain circuit 
model into a form in which the circuit can be analyzed us- 
ing only complex algebra. Hence the expressions of equa- 
tion 2.114 are called impedance transforms, and the 
transformed mathematical model is then in the impedance 
(or joj) domain. As a result, no differential equations are 
used to solve a steady ac circuit. 

All previous fundamental theorems, laws, and circuit- 
analysis techniques also apply to steady ac circuit analysis 
using impedances. Thus, an ac circuit representation in the 
impedance domain is analogous to a dc circuit model. On 
the other hand, the concept of impedance has no meaning 
in the time domain with time-varying voltages and currents. 

To demonstrate these concepts, consider the simple RL 
circuit in figure 2.74A, now with a complex voltage source, 
that is, a steady-state sinusoid defined as a phasor. Here, 
the current through the resistance and inductance is the 
phasor, I; therefore, 

v R = Tr, 

V £ ^Tjo)L. 



VI 



I v 



90° 



Pure resistance : 
V and I in phase 



Pure inductance: 
I lags V by 90° 



90° 



Vt 

Pure capacitance: 
I leads V by 90° 



Figure 2.73.— Voltage-current phasor relationships for cir- 
cuit elements. 



Impedance Transforms 

The current-voltage relationships for the three fun- 
damental elements have been found using phasors, as 



V = RI, 



V = jcuLI, 



V = 



I 

jo>C - 



These can be rewritten as voltage-phasor to current-phasor 
ratios: 



y-R, 



V • T 



I j«C" 



Vr V l 



V L =jIX L 




Vr I 



B 



Figure 2.74.— Steady sinusoid analysis of simple RL series 
circuit. 



54 



By KirchhofTs voltage law, 

V = V R + V, 



or 



V = IR + IjwL = I(R + jcoL). 



The impedance (equivalent) of the entire circuit is then 



»-? 



R + jcoL. 



(2.115) 



Because impedance is a complex quantity, it also has a polar 
form: 

Z = |Z| If, (2.116) 

where |Z| = (R 2 + (wL) 2 )^ mag nitude of impedance, Q, 

and = tan-H— ). 
R 

A phasor diagram for the circuit current and voltages is 
given in figure 2.74B. Note that as current is common to 
both elements, it could be used as the reference phasor. 
Here, voltage across the resistor, V R , is in phase with cur- 
rent, while that across the inductor, V £ , leads current by 
90°. The total circuit voltage, V, can be resolved noting that 

v = v* + v L = v |e, 

where V = ( \ R + V £ ) ^ magnitude of source voltage, V, 
and = tan-K— ). 

V 

This last angle is identical to that found for the impedance. 
It should be noted that the current and voltage relation- 
ships for the inductor are as those found previously when 
time-domain voltages and currents were considered. 

Now consider figure 2.75A, which shows a simple RC 
series circuit in which 

V* = IR, 

v c =T(^-)=-j-L, 



V= IR--^= I(R - -l_). 



and V = IR - -$- 

The impedance becomes 



T ojC 



R + 



jcoC 



(2.117) 



Figure 2.755, the circuit phasor diagram, shows the current- 
voltage phase-angle relationships with the voltage across 
the capacitor now lagging that across the resistor. 

Continuing the process for an RLC series circuit (fig. 
2.76A), the voltage across each element is 

V R = TR. 
V, = IjwL, 




v c 



V R =IR I 



Vc=-jix<;t 




Figure 2.75.— Steady sinusoid analysis of simple RC series 
circuit. 



j V R \^_ V C 



-VWW^ou — 1( — | 
■: R L C 





A B 

Figure 2.76.— Steady sinusoid analysis of simple RLC series 
circuit. 



and across the entire circuit, 

V = % + V L + v c 



or 



V = IR + IjwL + Kt^tt), 
jcoC 



(2.118) 



with the circuit impedance, 



or 



Z = R + jcoL + 



Z = R + j(coL 



jo)C 



o>C 



(2.119) 



The foregoing gives the essence of impedance transforms. 
Each impedance shown in equations 2.115, 2.117, and 2.119 
is the equivalent impedance of that circuit and has the 
general form 

Z = R + jX = |Z| \9, 

where R = resistance component, Q, 

X = reactance component, £2, 

|Z| = (R 2 + X 2 ) % , magnitude of impedance, Q, 

and = tan'HX/R). 

Here, depending on the pure circuit elements, the reactive 
component is 

X = ojL = inductive reactance, Q, 

X = — — = capacitive reactance, Q, 

o)C 

X = ojL = reactance for series LC elements, Q. 

o>C 



55 



From this equation, it can be seen that resistance is con- 
stant while reactance is variable with frequency. 

The time-domain expression found for a general series 
RLC circuit can be used to clarify the transformation 
process: 



v = iR + L i + i / .' idt + v - 



(2.93) 



It has been demonstrated in the impedance domain for 
steady ac that 



V = TR + TjcoL + T— — . 
jcoC 



(2.118) 



Accordingly, time-domain differential equations can be 
changed to the impedance domain when the circuit is under 
steady ac by 

1. Replacing v with V (in rms), 

2. Replacing i with I (in rms), 

j 

3. Replacing — with jco, 

dt 

4. Replacing / . . . dt with — , and 

5. Letting V = 0. 

However, it is a much more efficient approach to ac circuit 
analysis to assign the impedances directly using equation 
2.114, and soon more will be stated regarding this. 

Admittance 

Admittance, which is given the symbol Y, is defined as 
the reciprocal of impedance, Z, and 



Y = l 

V 



(2.120) 



The units are now Siemens, replacing the previous designa- 
tion, mhos. Admittance is therefore a complex quantity, the 
real part being conductance, G, and the imaginary compo- 
nent susceptance, B, or 



Y = G + jB. 



(2.121a) 



It should be noted that conductance is not the reciprocal 
of resistance unless reactance is zero, likewise for suscep- 
tance, reactance, and resistance. In general form, through 
equating Y and Z, 



R 



R 2 + X 2 



and B = 



-X 



R 2 + X 2 



(2.1216) 



Admittance affords basically the same convenience in 
steady ac circuit analysis that conductance provides for 
parallel dc circuits. 



Steady-State Analysis 

As previously stated, all circuit-analysis techniques that 
were covered for dc circuits still apply to steady ac circuits 
in the impedance domain. These include network reduction, 
Kirchhoff s laws, loop and node^analysis, network theorem, 
plus delta-wye transforms. Impedances simply replace 
resistances in the concept, and steady ac sources replace 
dc. Even with dc, the impedance domain can be used; in 
other words, dc sources can be thought of as steady-state 
sinusoids with w = 0. Therefore, with dc, reactance has no 
effect. 

A summary of circuit relationships follows, this time 
including impedance. 



1. Impedances in series. A single equivalent impedance, 



Z, is 



Z = Z x + Z 2 + Z 3 + + Z„. (2.122) 



2. Impedances in parallel. A single equivalent here is 

(2.123) 



1 = JL + JL + JL + .... + .L 

/j Zi} Zi 2 Zi 3 Zj„ 



3. Admittances in parallel, 

Y = Y, + Y 2 + Y 3 + . . . Y„. 



(2.124) 



4. Voltage distribution of series impedances, 

(2.125) 
where V is the input voltage, V x is across Z„ and so on. 



ZJi Zin 

V = —V V = —V 
V x Z V, V 2 z V, 



5. Current distribution through parallel admittances, 

(2.126) 



_ Y,- - Y 2 _ 
I 1 = Y I ' l2 = Y 1 ' 



where I is the total circuit current, I x is through Y„ and 
so on. Or parallel impedances, 



Z_ - Z _ 



(2.127) 



The overlines are removed on the above impedances and 
admittances simply for convenience, but it should be 
remembered that all are complex numbers. In essence, ac 
circuits in the steady state can be solved almost as easily 
as dc circuits employing only resistance. The major addi- 
tion is that the solution now uses complex algebra. 



56 



EXAMPLE 2.21 

Consider the circuit shown in figure 2.77, where 

v = 5,880 cos (377t + 53.1°), 
i = 141.4 cos 377 t. 

The circuit is under steady-state conditions. What are 
the values of R and L? 

SOLUTION. The phasor representations for voltage 
and current are 

V = ^HP 1 53.1° = 4,158 1 53.1° V, 

VT 



I= 14JL4 |()C 
\T2 - 



100|0° A. 



The total impedance of the circuit is then 

= 41.58 153.1° Q 



V = 4,158 |53.1' 

T 



100 |(P 

or in rectangular form, 

Z = 25 + J33.25 Q. 

The real part of this impedance must be the circuit 
resistance and the imaginary part equal to total reac- 
tance. Thus, 

R = 25 Q, 
X = 33.25 Q, 
but X = coL - 0.3. 



Therefore, as a> = 377 rad/s, 
33.25 + 0.3 



L = 



377 



0.09 H. 



— vwv 



vtO 



L 



^h 



-jO.3 

Figure 2.77.— Circuit for example 2.21. 



EXAMPLE 2.22 

Find the voltage, V, across the 2-Q resistance in 
figure 2.78. 

SOLUTION. Circuit reduction appears to be the 
easiest way to solve the problem. Noting that u> = 377 
rad/s, the reactances of the impedance and 
capacitance are 

X L = coL = (377)(0.12 x 10- 3 ) = 0.045 Q, 

1 1 



Xc = 



coC (377) (3,535 x 10" 6 ) 



= 0.75 Q. 



The impedance of the branch containing the im- 
pedance is 

Z t = R x + jXc = 1 + jO.045 Q 

for the branch with the capacitance 

Z 2 = R 2 + jX c = 1 - J0.75 Q. 

Combining these two parallel impedances in polar 
form, 



Z X Z 2 (1.0 1 2.58 "X 1.25 1 -36.87°) 



Z x + Z 2 



2-J0.705 



0.59 1-14.9° Q, 



and the equivalent impedance seen by the 1,000-V 
source is 

Z ea = Z + 0.59 1-14.9° 



= Z + 0.57 - jO.15 

= 2.57 - J0.5 = 2.57 1 -3.4° Q. 

Assigning the source voltage as the reference phasor, 
the total circuit current is 

- V = 1,000 |01 = 388 | 3 . 4 o A 
Z„ 2.57 1 -3.4° 

The voltage across the 2-S2 resistance is then 

V = 21 = (2)(388|3.4°) = 777|3.4° V. 



V 



l — vwv 

i 2n 

1,000 Vf A 
60Hz \ty 




Figure 2.78.— Circuit for example 2.22. 



57 



EXAMPLE 2.23 

Calculate the current, I, through the branch indi- 
cated in figure 2.79 using only loop equations. 

SOLUTION. Two loop currents have been assigned 
in the figure. Using Kirchhoffs voltage law, 

(5 - j5)I x - 5I 2 = 1,000|(T, 
-5I X + (5 + j5)I 2 = 800|0°_. 

The solution to these simultaneous equations gives 

l x = 360 + j200 A, 
I 2 = 360 - jl60 A, 

The current through the 5-Q resistance is then 

I = Ix - I 2 

or I = 360 + J200 - 360 + J160; 

thus, I = j360 A 

or I = 360190° A. 



EXAMPLE 2.24 

What are the Thevenin's and Norton's equivalents 
for the circuit shown in figure 2.80? 

SOLUTION. Applying either Thevenin's or Norton's 
theorem, the equivalent impedance of the circuit be- 
tween a and b with the internal source off is Z s . When 
the steady-state voltage source is off, it acts as a short 
circuit, and the 4-Q and 12-Q resistances are effec- 
tively in parallel, and 



(4) (12) 
4+12 



3 Q. 



This combined resistance is in series with the jlO-Q 
reactance, and the series combination is in parallel 
with the — j6-Q capacitance, and 

z = z (3 + jl0)(-j6) . 

° b 3 + jlO - j6 ' 
thus, Z, = 12.5| -69.8° = 4.3 - jll.7 2. 

If the terminals a and b are shorted out according to 
Norton's theorem, a short circuit exists across the 
capacitance, and the jlO-Q impedance and 12-Q resist- 
ance are placed in parallel. The equivalent impedance 
of the circuit under this shorted condition as seen by 
the ideal voltage source is then 

Z„ = 4 + (12)( J 10) = 10.7| 33.5° Q. 
12 + jlO 

The circuit delivered by the ideal source is 

V 25|0 C 



Ix = 



10.7133.5' 



= 2.33-33.5 A, 



and the current through the jlO-Q reactance and the 
shorted terminals is 



I 2 = I„ 

or I, = 2.33 1 33.5 (- 



I = Ix L„ 12 .., 



12 



15.62 39.8° 



12 + jlO 

) = 1.8|-73.3° A. 



Z s and I s define the components of the Norton 
equivalent for the circuit in figure 2.80. By source 
transformation, 

V s = I S Z S = (1.8| -73.3° )(12.5l -69.8° ) 
= 22.5| -143.1° V, 

which is identical to 

V* = -22.51 36.9° V. 

Z s and V s define the components of the Thevenin 
equivalent. 

It can be noted in figure 2.80 that the ideal volt- 
age source is in series with the 4-Q resistance in one 
branch. Therefore, source transformation alone could 
be employed to solve the problem. 

The use of subscripts in this example did not 
follow the format previously used in the chapter. In 
other words, Z„ I„ and V £ described the equivalent 
circuits rather than Z , I , and V . The reason is 
that the subscript zero has a special meaning in 
three-phase ac circuits, which will be discussed in 
chapter 4. 



j5n 



or vf©2) iji 5 n (V ©| 80 °I0-° V 
I— H( * 1 

Figure 2.79.— Two-loop circuit for example 2.23. 



4fL 



jion 



25 




loivf© 



Figure 2.80.— Active circuit for example 2.24. 



58 



Chapter 2 has introduced the concepts of electrical cepts are fundamental to electrical engineering, regardless 

circuit analysis. The fundamental laws were covered first, of application. Thus, comprehension of the contents of this 

followed by numerous circuit analysis techniques, which chapter is vital to understanding the following chapters, 

were applied to dc circuits. Steady ac was then presented, The next chapter will continue the study of electrical fun- 

and the chapter concluded with examples of circuit analy- damentals, with emphasis on power consumption in ac 

sis on ac circuits under steady-state conditions. These con- circuits. 



59 



CHAPTER 3.— ELECTRICAL FUNDAMENTALS II 



The measures of instantaneous power, p, and average 
power, P, were introduced in chapter 2. Instantaneous 
power does not have application in steady ac circuit 
analysis, so the concept of average power has been devel- 
oped to gauge the rate at which electricity does work. This 
chapter continues to build the foundations for mine power 
fundamentals that will be expanded into full comprehen- 
sion in chapter 4. There, the discussion will focus on 
three-phase power; here, the purpose is to introduce single- 
phase power and transformers. 



AVERAGE POWER AND POWER FACTOR 

To find the average power consumed by a circuit, the 
resistance of each element can be examined and all the 
individual power consumptions computed. Reactance, ei- 
ther capacitive or inductive, does not affect average power. 
When all the average powers have been determined, their 
sum yields the total average power delivered to the circuit. 
Obviously, if the circuit elements are numerous, the pro- 
cess can be time consuming, but this approach is some- 
times necessary. 

If the average power needs to be determined for the 
total circuit, it would be more desirable to perform only 
one calculation by computing average power in terms of 
the terminal current and voltage in the circuit. Yet, when 
complex or imaginary components exist in the circuit, can 
they be ignored, as this implies? In other words, the 
voltage and current waveforms might not be in phase, and 
when a phase angle is involved, the product of effective 
voltage and current no longer equals average power. 
However, instantaneous voltage and current can be used to 
calculate average power and to demonstrate what occurs if 
a circuit has reactance. 

Assume that the following current and voltage are 
monitored at the terminals of a circuit: 

i = ImCOS wt, 

v = V m cos(a>t + 0). 

Current is taken as reference, and the phase angle by 
which voltage leads current is 0. The instantaneous power 
consumed is then 

p = vi = V m I m cos(wt + 0)cos tot. 

From the trigonometric identity for the product of two 
cosines, 



V L. 
p = -son (CO80 + cos(2wt + 0)) 



The first term of equation 3.1 is constant, while the second 
is a sinusoid. Thus, taking the average to find average 
power results in 



V I 

P = av(p) = -^ cos 0. 



(3.2) 



Realizing that V„ 
becomes 



= V2V and I m = V2I, average power 



P = VI cos 



(3.3) 



or 



P = 



V^ 



cos + 



V„Xn 



co8(2wt + 0). 



(3.1) 



in which V and I are root-mean-square (rms) voltage and 
current at the circuit terminals and is their phase angle. 
If the voltage and current had been dc values, the average 
power would just be the product of voltage and current. 
However, when voltage and current are sinusoidal, equa- 
tion 3.3 specifies that the average power entering any 
circuit is the product of the effective voltage, effective 
current, and the cosine of the phase angle. 

The function cos is called the power factor (pf). For a 
purely resistive load, the phase angle is zero and the power 
factor is unity. Unity power factor may also exist when 
inductance and capacitance are present, if the effects of 
reactive elements cancel. If the circuit is totally reactive 
(either inductive or capacitive), the phase angle is a 
positive or negative 90°, the power factor is zero, and 
average power must be zero. 



COMPLEX AND APPARENT POWER 

When there is reactance in a circuit, a component of 
circuit current is used to transfer stored energy. The 
energy is periodically stored in and discharged from the 
reactance. This stored energy adds to circuit current but 
not to average power because average power to reactive 
elements is zero. In such cases, the power factor is not 
unity. Thus, as no work is performed by the added current, 
the power factor can be considered to be a measure of 
circuit efficiency or its ability to perform work, and aver- 
age power, defined by equation 3.3, is often called active 
power or real power. 

Power calculations can be simplified if power is de- 
fined by the complex quantity shown in figure 3.1, which 
is expressed mathematically as 

S = P + jQ, (3.4) 

where S = complex power, 

P = real power, as before, 
and Q = reactive power or imaginary power. 

Imaginary power accounts for the energy supplied to the 
reactive elements. If 

P = VI cos0, 



60 














s 


i 


1 




>■ 






ir 






< 






-z. 






o 






< 






^ 


/\ e 


I 

i 




REAL 


P 



Figure 3.1. —Power represented as real and imaginary com- 
ponents. 



then the magnitude of complex power, S, called apparent 
power, is 



S = VI, 
and imaginary power is 

Q = VI sine. 



(3.5) 



(3.6) 



Voltage and current are again rms, and is the phase 
angle. Therefore, 



or 



S = P + jQ = VI cos0 + jVI sine 
S = VI(cos0 + jsin0) = Vie* = VI|0. 



(3.7) 



Complex power is then simply the product of terminal rms 
voltage and current magnitude acting at a phase angle. 
Applying dc concepts, the product, VI, is the power appar- 
ently absorbed by the circuit, hence the term apparent power. 
Apparent power, real power, and imaginary power are di- 
mensionally the same, but to avoid confusion with real 
power (units of watts), apparent power has units of voltam- 
peres, and reactive power uses voltamperes reactive. 

When sinusoidal voltage and current have general 
form, as in 

V = V|0, 

I = I|0, 

instead of using equations 3.4 and 3.7, the following 
expression is more convenient for computing complex 
power: 



S = VI*, 

where V = complex voltage, V, 

and I* = conjugate of complex current, A. 



(3.8) 



Accordingly, 



or S = Ve^e*-* = Vie**-**, 

where - 4> = phase angle between voltage and current. 



EXAMPLE 3.1 

When operating under normal conditions, an 
induction motor has been found to draw 100 A when 
440 V is across its terminals. Current is lagging 
voltage by 36.87°. Find the average, reactive, appar- 
ent, and complex powers for this load. 

SOLUTION. From equation 3.3, the average power 
is 

P = (440X100) cos 36.87° = 35,200 W. 

Using equation 3.6, the reactive power is 

Q = (440X100) sin 36.87° = 26,400 var. 

Equation 3.5 defines the apparent power as 

S = (440X100) = 44,000 VA, 

and equation 3.4 yields the complex power as 

S = 35,200 + j26,400 VA. 

ALTERNATIVE SOLUTION. If voltage is assigned 
as the reference phasor, then 

V = 440|0°^ V, 

I = 100| -36.87° A. 

From equation 3.8, the complex power is 
S = (440|0°X100| -36.87° )* 

or S = (440|0°X1001 36.87° ) = 44,0001 36.87° VA, 

where the magnitude is the apparent power, or 

S = 44,000 VA. 

Converting the polar expression for complex power 
to a rectangular form, 



S - VI* = V\0I\-4> = VIl0-4> 



which yields 



S = 35,200 + j26,400 VA, 

P = 35,200 W, 

Q = 26,400 var. 



It should be noted that the above solutions are 
only two of the many possible. 



61 



EXAMPLE 3.2 

A load consumes 1,250 kW at 0.6 lagging power 
factor when 4,160 V at 60 Hz is across it. The load is 
connected in series with a (0.71 + jO.71) fl impedance 
to a constant source. Determine the voltage and 
power factor at the source. 

SOLUTION. From the stated conditions, the aver- 
age power is 

P 1 = 1,250 kW. 

From equation 3.3, the current through the load is 

Pi 



Ii = 



V x cosd 1 ' 



where P x , V 1( and cos^! relate the conditions for the 
load, or 

I _ 1 > 250 > Q0Q _ sol A 
x i - (4,160) (0.6) " 0U1 A - 

For convenience, the voltage across the load can be 
assigned as the reference phasor, then 

V\ = 4,160 V, 

I = 501|^L = 5011 -53.1° A. 

The load current also flows through the series im- 
pedance. Using polar expressions, the voltage drop 
across this impedance is 

V 2 = I X Z 2 

or V, = (501| -53.1 o Xl|45°) = 501| -8.1° V. 

The voltage at the source is then 

v 8 = V\ + v 2 

= 4,160 + (496 - j71) 

= 4,656 - j71 = 4,6571 -0.9° V. 

The power factor at the source can be found by first 
calculating the phase angle between current and 
voltage at the source with the current phasor taken 
as reference. Here, 

8 = <?v _ #1 
or 6 B = -0.9° - (-53.1°) = 52.2°. 

Therefore, the power factor at the source is 
cos 6 e = cos 52.2° = 0.61 lagging. 



Any circuit in steady state can be reduced to the general 
impedance 



or 



Z = R + jX 



Z = |Z| \§_ = |Z|e* = j e* 



(3.9) 



where 



= tan 



JR. 



By relating the complex power consumed by a circuit with 
this impedance, another useful expression can be found: 



hence, 



S = VI e* = I 2 (y e* j = I 2 Z; 
S = I 2 (R + jX) = I 2 R + jI 2 X. 



Because S = P + jQ, 

P = I 2 R, Q = I 2 X 
or P = V R 2 /R, Q = V x 2 /X. 



(3.10a) 
(3.106) 



In equation 3.10a, the current I is real rms, not complex. 
The rms voltages in equation 3.106 are those existing 
across the individual elements, not across the total circuit. 
It may already be obvious that the circuit impedance 
angle is identical to the power-factor angle. The following 
also apply. 

1. If a circuit contains resistance and capacitance (a 
capacitive load, Z = R - jX), the current leads voltage, the 
phase angle is negative, and Q is negative. 

2. If the circuit is an inductive load (Z = R + jX), the 
current lags voltage, and the phase angle and Q are 
positive. 

In either case, the power factor ranges from zero to unity 
(purely reactive to purely resistive). A capacitive load is 
said to have a leading power factor, an inductive load a 
lagging power factor, as illustrated in figure 3.2. 



P=VIcos 




0=VIsin8 
Leading 




Q=VIsin( 
Lagging 



A Capacitive load 



P=VIcos 
B Inductive load 



Figure 3.2.— Illustration of leading (A) and lagging (5) power 
factors. 



62 



The complex power delivered to several loads is the 
sum of the complex power consumed by each individual 
load, no matter how they are interconnected. This rela- 
tionship can be shown using the simple circuit in figure 
3.3. The total complex power to the system is 



but 
thus, 



S = V I* = P + jQ, 
1 = 1! + I 2 ; 

S = V(I* + I£) 

= v if + v if = s x + s 2 



or 



S = P, + F 



2 + JQi + jQ 2 - 



This has extensive practical significance. For example, if a 
circuit has a lagging power factor, a capacitance (with a 
leading power factor) can be selected and then placed in 
parallel, so as to negate or reduce the total circuit imagi- 
nary power (with the capacitance). The net result is to 
reduce total circuit current, while the load consumes the 
same real power and thus performs the same work. This is 
the essence of power-factor improvement. 



EXAMPLE 3.3 

Consider that the two loads shown in figure 3.3 
are induction motors operating as follows: 

P x = 50 kW at 0.6 lagging power factor, 
P 2 = 25 kW at 0.8 lagging power factor. 
Find the overall apparent power and power factor 
when these consumptions are combined. 

SOLUTION. The average, apparent, and reactive 
power for each load are 



P 1 = 50 kW, 



Si- 



P 1 50,000 



= 83.33 kVA, 



cosd 1 0.6 
Q x = S x sinfli = 83,333(0.8) = 66.67 kvar, 



P 2 = 25 kw, 



B.--&- 



25,000 



= 31.25 kVA, 



cos0 2 0.8 
Q 2 = S 2 sin0 2 = 31,250(0.6) = 18.75 kvar. 

Complex power is then 

S = P x + P 2 + jQ x + jQ 2 

or S = 50 + 25 + J66.67 + J18.75 

= 75 + J85.42 kVA. 

Apparent power is the magnitude of complex power, 
or 

S = (P 2 + Q 2 ) 1/2 

= (75 2 + 85.42 2 ) 1/2 = 113.7 kVA. 



o 

T 


— < 


1 1 


I* 


V 


Pi 




P2 





— « 


> 





Figure 3.3.— Circuit demonstrating sum of complex powers. 



The power-factor angle is 
= Tan- 1 ^ 



(P 

! /85.42 



- Tan ' ( :: ^) = 48.72°, 



and the power factor of the combination is 
pf = cos = cos 48.72° = 0.66. 



EXAMPLE 3.4 

The maximum capacity of a piece of power equip- 
ment is rated by apparent power at 500 kVA. The unit is 
being loaded by 300 kW at 0.6 lagging power factor. The 
power factor must be improved to 0.8 lagging by adding 
capacitance in parallel with the equipment. Find the 
required capacitance in kilovoltamperes reactive. With 
the capacitance in place, find the reserve capacity that is 
available from the power equipment. 

SOLUTION. For the load on the equipment without the 
capacitance, 



P, = 300 kW, 



Pi 



300 



= 500 kVA, 



1 cos0, 0.6 
Q x = SiSinflx = 500(0.8) = 400 kvar. 

It can be said from S x that the equipment is fully 
loaded. When pure capacitance is added, average 
power remains constant, and only reactive power 
and apparent power change. For the desired power 
factor, cos0 2 , 



EL- 



gi 

COS0Q 



300 
0.8 



= 375 kVA, 



Q 2 = S 2 sin0 2 = 375(0.6) = 225 kvar. 



63 



Consequently, the added capacitance causes the 
total reactive power to decrease. The difference 
between the reactive power without and with the 
capacitance must be the amount inserted by the 
capacitance. In other words, 

Q c = -<Qi - Q 2 )> 

Q = -(400 - 225) = -175kvar. 

The negative sign is used here to indicate that the 
capacitance adds negative reactive power to the 
system. Finally, the difference between the apparent 
power without and with the capacitance yields the 
reserve capacity available from the equipment, or 

= 500 - 375 = 125 kVA. 

It can be noted that additional average power can 
now be added to load on the equipment without 
exceeding its maximum capacity. For instance, con- 
sider that average power P will load the equipment 
so that the equipment is again operating at full 
capacity. Then, the total average power is 

P T = 300 kW + P. 

Reactive remains constant, 

Q T = Q 2 = 225 kvar, 

and apparent power changes to 

St = 500 kVA. 

Therefore, St = (P T 2 + Q T 2 ) 1/2 , 

500 = [(300 + Pf + 225 2 ] 1/2 . 

Solving for the new average power, 

P = 146.5 kW. 



At w , the circuit is said to be in resonance, and 



2 1 1 

u ° = LC 0r '° o = (LC7 75 - 



Since a) = 2irf, the resonance frequency, f , is given by 

1 



f„ = 



21KLC) 1 



(3.13) 



For a series RLC circuit in resonance, it can be shown that 

1. The applied voltage, V, and the resulting current, I, 
are in phase, 

2. The power factor of the circuit is unity, 

3. The impedance, Z, is minimum, and 

4. The current, I, is maximum. 

At all other frequencies that are significantly higher or 
lower than f , the series RLC circuit appears as a high 
impedance. With frequencies below resonance, capacitive 
reactance is greater than inductive reactance, so the angle 
of impedance is negative (total reactance is negative). 
Above resonance, the situation reverses and the imped- 
ance angle is positive. This can be seen clearly in figure 
3.5 where circuit impedance versus frequency is plotted. 

The energy stored in a resonance circuit is essentially 
constant, yet the energy level within the circuit may be 
many times higher than the energy being supplied from 
an external source during any period. The source itself 
does not supply any reactive power, only active power. The 
reactive power transfers energy back and forth between 
the resonant-circuit inductance and capacitance. The re- 
sult of this energy transferral can be very high voltages, 
several times the terminal voltage, existing across the 
inductance and capacitance within the resonant circuit. 



o— WV^TP — 1( — o 
RLC 

Figure 3.4.— Simple series RLC circuit for resonance. 



RESONANCE 
Series Resonance 

Earlier, the impedance for the simple series RLC 
circuit shown in figure 3.4 was found to be 

Z = R + jcoL + j-q = R + j( u L - ^ ). (2.119) 

A special circuit phenomenon can now be demonstrated 
with this equation. There exists one specific frequency, w , 
where total circuit reactance is zero and the circuit imped- 
ance is purely resistive, or 



w„C 



w„L - — = 0or w L = — p (3.11) 



1_ 

w„C 



UJ 

o 
< 

Q 
LU 
Q. 




CD 



R 






x L 


= ojL 




Z = 


t/r 2 h 


■X2 


x = 


ojL- 


J_ 


x= 


1 





and 



Z„ = R. 



(3.12) 



FREQUENCY (co), rad/s 

Figure 3.5.— Plot of impedance magnitude versus frequency 
for series RLC illustrating resonance. 



64 



This situation can be the cause of some severe overvoltages 
in mine power systems, and the concept will be explored 
further in chapter 11. 

The amount of energy stored, compared with that 
dissipated by the resistance, is related to the shape of the 
curve representing impedance magnitude, as shown in 
figure 3.5. This curve is an example of a response curve. 
The quality factor of a circuit is a measure of the sharp- 
ness of the response curve and is expressed as a ratio: 



Q„ = 2* 



maximum energy stored per period 
total energy lost per period 



(3.14) 



where the period is one complete cycle of the resonant 
frequency. By finding the ratio of the energy stored in 
either of the circuit's reactive components to the energy 
dissipated in the resistance, it can be shown that 



Q = 



reactance 
resistance 



R 



„CR 



(3.15) 



The quality factor normally has greater application in the 
communications aspects of electrical engineering than in 
the power aspects. For instance, the width of the response 
curve is also related to Q and has great relevance to the 
tuned circuits used in radio and television. 

Parallel Resonance 

The resonance of the simple parallel RLC circuit 
shown in figure 3.6A is very similar to that just discussed. 
This circuit is obviously idealized, but its performance is of 
general interest. The admittance can be written as 

Y = G + juC + j^l = G + j( w C - ^ ), (3.16) 

and the circuit is in resonance when susceptance B is zero. 
Hence, the circuit exhibits low admittance and high 
impedance at resonance, while the series RLC circuit had 
low impedance and high admittance: 



B = a. C - — j- = or co C = — j- 



(3.17a) 

or Y = G. (3.176) 

On the other hand, the resonant frequency is again 

1 



f„ = 



Therefore, Q of this parallel resonant circuit is the dual of 
equation 3.15 or 



Q = 



susceptance « C R 
conductance G «„L 



(3.18) 



The concept also relates to many fundamentals covered in 
chapter 2. For example, two circuits are called duals if the 
loop equations for one have the same forms as the node 
equations for the other. 

Because figure 3.6A is idealized (as actual inducting 
elements must have associated resistance), figures 3.6S 
and 3.6C are presented to show practical circuits that 
exhibit parallel resonance. 



TRANSFORMERS 

Early in chapter 2, the concept of mutual inductance 
was introduced. To review, Faraday found that a time- 
varying current in one circuit would induce a voltage in a 
nearby circuit. If the adjacent circuits are simply conduc- 
tors and are labeled 1 and 2, as in figure 3.7, this 
statement means that 

• i x in circuit 1 produces v 2 in circuit 2, 

• v 2 in turn causes i 2 to flow (if circuit 2 is part of a 
complete loop), then 

• i 2 induces v x in circuit 1. 

These interrelated phenomena can be thought of as mag- 
netic coupling between the two circuits, and it has been 
shown that 



di 2 di x 

= L 12 — andv 2 = L 21 — 



where L 12 = L 21 = M = mutual inductance, H. 



C£ 



C~ 



27KLC) 1 



The statements previously given for series circuits also 
apply, except that current replaces voltage and voltage 
replaces current. 

This is an example of duality. Anything stated about 
a series resonant circuit applies to its dual, the parallel 
resonant circuit, if each word in the left column below is 
replaced by its opposite word shown in the right column: 

Series Parallel 

Voltage Current. 

Impedance Admittance. 

Resistance Conductance. 

Reactance Susceptance. 

Inductance Capacitance. 



Figure 3.6.— Circuits that exhibit parallel resonance. 



Flow of current 
causes magnetic 
field that cuts 
other conductor 



U fet 



V, 




Magnetic flux lines 
Figure 3.7.— Magnetic coupling between two conductors. 



65 



Because of the equality, M is used to represent mutual 
inductance. These equations are true only for straight 
wires, and magnetic coupling exists only if voltage and 
current are time varying. 

The circuits considered previously were loops or 
meshes composed of passive and active elements, and 
these were conductively coupled by common branches or 
nodes. The following paragraphs develop the concept of 
magnetic coupling further and introduce the fundamen- 
tals behind one of the more important components of ac 
mine power systems, the transformer. 

Transformers are prime examples of magnetic cou- 
pling. They are often designed to optimize this coupling, 
and their operation is based inherently on mutual induc- 
tance. Transformers are employed to increase the magni- 
tude of voltage for more economical power transmission or, 
conversely, to decrease the level to provide voltage more 
suitable for electrical equipment operation. In essence, 
these changes can be made with either total isolation or 
direct conduction between circuits. 

Instead of straight conductors, assume that two coils 
are situated side by side, and their magnetic action is 
passing through any environment (fig. 3.8). The current in 
coil 1 is then partly the result of self-inductance in coil 1 
and mutual inductance from coil 2, and vice versa for coil 
2. Expressed mathematically: 



_ d^ _ , di 2 



dii , dio 

= <iM dT + L *df 



(3.19a) 



(3.196) 



where L x , L 2 = self-inductances of coil 1 and coil 2, respec- 
tively, H, 
and M = mutual inductance, H. 

The additional terms implied by these equations exist only 
if more than two coils (or circuits, or windings) are 
interacting, and they are presented merely to make the 
expressions more general. 

The plus and minus terms of the equations deserve 
special attention. Sign convention has been well defined 
for inductors, and coil 1 and coil 2 are inductors when 
taken individually. A current flowing into the coil pro- 
duces an opposing voltage, hence the positive sign or 
polarity. The potential created by mutual inductance, M, 
however, cannot be treated in the same manner. This 
voltage may have either positive or negative polarity 
depending on the winding sense, the direction the coils are 
wound with respect to one another. Consider the two coils 
wound on a common core in figure 3.9A. They are wound 
in the same direction and therefore have the same sense. If 
a current is flowing into the top of the upper coil, the 
voltage produced by this current adds to that produced by 
the same current direction in the lower coil. But in figure 
3.9B, the winding sense of the lower coil is reversed so that 
the same current in the top coil now creates a voltage that 
opposes the current produced in the lower coil. Therefore, 
the polarity of mutual-inductance voltages can be found by 
drawing physical sketches. However, this is impractical in 
circuit diagrams, and so magnetically coupled coils are 
often marked with dots that represent the direction of 
polarity. A dot is placed at the terminals of the coils that 



are instantaneously at the same polarity as a result of 
mutual inductance. Thus, in figure 3.10A, i 1 enters the 
dotted terminal of L x , v 2 is sensed positively at the dotted 
terminal of L 2 , and 



M f' 



dt 



_ , di, T di 2 

V2 = M dT + L2 df 



(3.20a) 



(3.206) 



In analyzing circuits, it may be more desirable to reference 
v 2 as positive at the undotted terminal of L 2 , as in figure 
3.10.B. In this case, 



. di, ,„di 2 



,,di, T di 2 

V2= - M dT + L2 df 



(3.20c) 



(3.20d) 



What is important is that, in either instance, the mutual 
voltage is produced independently from that of self- 
induction. 



^-Magnetic coupling 




h — — <"' ' 




— II 


^ j 


v 2 


Load 




— ii 



L ^S^'J L 2 



M 



Figure 3.8.— Magnetic coupling between two coils. 



o > 

« — o 



o =- 

i : 




) 






< 




, Z 

o 

• 



A B 

Figure 3.9.— Demonstration of coil winding sense. 



Ml Li N 2 L 2 _ i, . • • i 2 

f N U|IC N 2 | v 2 Polarity 

V,[ L 5 t|_ | v 2 change is 

'[ I 2 for equation 
M 3.20 only 

A Actual winding sense B Dot notation 

Figure 3.10.— Dot convention for mutual inductance sign. 




66 



The equations just presented are valid for any voltage 
or current waveform. If the currents are sinusoidal and 
have a radian frequency, w, transforms can be employed so 
that for equations 3.20c and 3.20d, 



Vi = juL^ - jcoMI 2 , 
V 2 = -jwMlj + jwL 2 I 2 . 



(3.21a) 
(3.216) 



These relationships can be used to analyze circuits con- 
taining magnetically coupled elements. It should be stated 
that equations 3.20 and 3.21 relate only to the magneti- 
cally coupled elements; equations for complete circuits 
containing these devices will follow. 



IDEAL TRANSFORMER 

The level of mutual inductance, M, depends upon the 
spacing and orientation of the coils and the permeability 
of the medium between them. In other words, M is a 
function of the magnetic flux linking between the coils. 
More will be said about this phenomenon later in the 
section. In figure 3.10A, by comparing the power entering 
L x of the circuit with that stored or available in L 2 , it can 
be proved from flux-linking concepts that 



M < (I^L,) 1 



(3.22) 



Consequently, M has an upper limit defined by the geo- 
metrical mean of the two inductances involved. The ratio 
of M to its theoretical maximum is called the coefficient of 
coupling. This is by definition 



k = 



M 



(L^) 1 



(3.23) 



where k can range from zero to unity. Coils having a low 
coefficient of coupling are said to be loosely coupled. Here 
the coils could be far apart or have an orientation such 
that little magnetic flux interacts between them. Loosely 
coupled circuits may have a k that ranges between 0.01 
and 0.10. For tightly coupled circuits, such as air-core 
coils, k can be around 0.5. 

A power transformer is a device having two or more 
tightly coupled coils or windings on a common iron core. 
The coils are wound and oriented to provide maximum 
common magnetic flux and can have a coefficient of 
coupling very close to 1.00. The usual range is 0.90 to 0.98. 
Resistance and other power losses are small. The winding 
receiving power is called a primary; that delivering power 
is called a secondary. In the circuit in figure 3.10, h 1 is the 
primary and L 2 is the secondary. An ideal transformer is 
an idealized form of transformer where k = 1 and losses 
within the device are zero. Hence, an ideal transformer 
can deliver all the power it receives. Many useful relation- 
ships for real transformers can be obtained by assuming 
the ideal transformer case. 

The self-inductance of a coil has been shown to be 
proportional to the square of the number of turns forming 
the coil (N), provided that all the flux, created by the 
current in the coil, links all the turns (see chapter 2, 
"Inductance"). If a sinusoidal current, I, flows in a coil of 
N turns, then the voltage produced across an N-turn coil 
must be N times that caused in a 1-turn coil. Further, for 



a sinusoidal voltage, V, which is constant across an N-turn 
coil, the current allowed through must be 1/N times that 
caused in a 1-turn coil. Both these statements can be 
proved by magnetic field concepts, again assuming that all 
magnetic flux produced in a coil links all turns. It follows 
that for an ideal transformer with two windings: 



L, 



Nf 
N5 



N, 



Vx N/ 



I 2 N x * 



(3.24) 



(3.25) 



(3.26) 



where 



N x = number of turns in primary winding, 
N 2 = number of turns in secondary winding, 
L 1; l lf V x = primary winding inductance, rms cur- 
rent, and rms voltage, respectively, 
and L 2 , I 2 , V 2 = secondary winding inductance, rms cur- 
rent, and rms voltage, respectively. 

For this two-winding arrangement, the voltage and cur- 
rent can be complex sinusoids. The turns ratio of the 
transformer, a, is defined as the ratio of the number of 
turns in the secondary winding to the turns in the primary 
winding: 



a = 



N2 
N x 



(3.27a) 



Hence, for an ideal transformer, 

W y, i 2 



(3.276) 



In other words, the sinusoidal voltages across the primary 
and secondary windings are in direct proportion to the 
number of turns of the windings, and the currents are 
related inversely to the turns. In addition, the last equa- 
tion shows that the apparent power at the primary and 
secondary windings is indeed equal: 

VA = v 2 i 2 . 

The magnitude of this power in voltamperes is specified 
for the maximum allowable or rated capacity of power 
transformers. 

Another useful transformer relationship can be deter- 
mined through a demonstration of steady ac circuit anal- 
ysis with magnetically coupled circuits. Consider figure 
3. HA, where a sinusoidal voltage source, V s , with an 
internal impedance, Z g (the combination is the Thevenin 
equivalent for a source), is connected to the primary of an 
ideal transformer. The secondary delivers power to a load 
impedance, Z L . The vertical lines between the transformer 
windings indicate that the core is made of iron lamina- 
tions. The turns ratio above the transformer symbol, l:a, 
relates a convention of N 2 to N x . 

A very useful relationship is the ideal-transformer 
input impedance with the load connected, that is, the load 



67 




Ideal transformer 




B 



Figure 3.11.— Demonstration of impedance transfer in 
transformers. 



or rearranging, 



7 _ 7 J^LiZ L 

^-^Z^jcoa 2 ^ 



Now allowing L x to tend toward infinity, the input imped- 
ance for the voltage source becomes 



Z in = Z B + -| 



(3.28) 



that the source sees through the transformer. Loop equa- 
tions can be used to solve the problem. Two loops, \ x and I 2 
(both express complex currents), are available in the 
circuit; the loops are magnetically coupled through the 
transformer. Employing Kirchhoff s voltage law for loop 1, 

V s = I^g + IJuL,! - LjuM, 

and for loop 2, 

O = -IjjwM + I 2 Z L + LjcoL. 

M is again the mutual inductance. Notice that current 
enters the dot of the primary and leaves the dot on the 
secondary, making the sign of M negative. Rewriting these 
into standard loop-equation form gives 

V s = IjfZg + j<oL x ) - y«M, 

O = -IjwM + I 2 (Z L + ja>L 2 ). 

Solving for I ls 



V 8 = IiCZ. + j«L x ) + U 



) 



Z L + jwL, 



Therefore, the impedance seen by the source, Z in , is the 
ratio of the source voltage to terminal current, or 

V w 2 M 2 

Ii Z L +jwL 2 



but 
then 



M 2 = L X L 2 



Z in = Z_ + jwLi + 



w L X L 2 
Z L + jcoL 2 



There must be total coupling between primary and sec- 
ondary windings for an ideal transformer; thus, the self- 
inductances, L x and L 2 , have no effect in the circuit, and 
their value can be considered infinite. Notwithstanding, 
the ratio is still finite, as specified by the turns ratio: 

L 2 = a L x . 

For this reason, primary and secondary inductances are 
conventionally not specified on ideal transformers. When 
this is related to the input impedance expression, 



Z in = Z g + jcoLi + 



co 2 a 2 L 1 2 
ZL+joa 2 !^ ' 



Equation 3.28 is significant as it shows that the 
source sees the load impedance, Z L , through the trans- 
former as Z L /a 2 . This means that an ideal transformer has 
the capability to change an impedance magnitude. There- 
fore, to assist in circuit analysis, the circuit in figure 3.11A 
can be redrawn to its equivalent, shown in figure 3.1 IS. 
Here, the impedance connected to the secondary is trans- 
formed to the primary. Obviously, the reverse process, 
primary to secondary, also holds, but the impedance is 
multiplied by a 2 . The impedance angle remains constant 
in either situation. 



EXAMPLE 3.5 

A 60-Hz single-phase transformer has a rated 
capacity of 250 kVA and a turns ratio of 15:1. 
Assuming that the transformer is ideal, find the 
primary voltage if the secondary voltage is 480 V. 
What are the magnitudes of primary and secondary 
currents with these voltages applied and the trans- 
former operating at full capacity? 

SOLUTION. For the turns ratio of 15:1, 



a = 



15 



As the turns ratio specifies the secondary voltage to 
the primary, 



a = 



Y2 

and the primary voltage is 
V 



Vj = — = 15(480) = 7,200 V. 
a 



The primary current for 250 kVA at 7,200 V is 



_ 250,000 
1 " 7,200 " 6 ° A ' 



and the secondary current for 250 kVA at 480 V is 



_ 250,000 _ 
12 " 480 " ° 



68 



EXAMPLE 3.6 

Consider that the circuit shown in figure 3. HA 
has the following parameters: 

Z g = 6 + j3 n, 

Z L = 1 + jO.5 «, 

V 8 = 7,200 V, 60 Hz, 

Turns ratio = 12:1. 

Find the value of the load impedance (Z L ) referred to 
the transformer primary, the complex power at the 
source, the transformer secondary voltage and cur- 
rent, and the required transformer capacity. 

SOLUTION. For the specified turns ratio, 

1 



a = 



12" 



Transferring the load impedance to the primary, 
\ = (12m + jO.5) = 144 + J72 «, 



which is the impedance referred to the transformer 
primary. The total impedance seen by the source is 
then 



Ze, = Z g + 



= 6 + j3 + 144 + j72 = 150 + j75 «. 

Assigning the source voltage as the reference pha- 
sor, the transformer primary current is 



Ii = 



7,200 1 0J 



= 42.91 -26.6° A. 



The transformer secondary current is 



I 2 = - = 12(42.91 -26.6° ) = 515| -26.6° A, 
a 



and the secondary voltage is 

V 2 = I 2 Z L = (515| -26.6° X1 + jO.5) 

or V 2 = (5151 -26.6° X1-12| 26.6° ) 

= 576|0^ V. 

The complex power delivered to the load is then 

S = V 2 I 2 * 

or S m (576|0°X5151 26.6° ) = 296| 26.6° kVA. 



This may also be found from 

S = I 2 Z L = (515m + jO.5) 
= 265 + J133 kVA 



or 



S = 296126.6° kVA. 



Finally, the apparent power demanded by this load 
is the required transformer capacity, 296 kVA. 



ACTUAL TRANSFORMERS 

In actual transformers, a source must furnish the 
power dissipated by the secondary load plus the power 
needed to operate the transformer. The additional power is 
created from losses within the transformer circuit. The 
transformer capacity, the amount of power it can handle, is 
dependent upon the character of these losses, which are 
dissipated as heat in the core and the windings. Because 
excessively high temperatures are destructive to insula- 
tion, the capacity is limited by this rise in temperature, 
usually specified as an allowable temperature rise above 
ambient conditions. 

The major losses in an iron-core transformer are 
winding resistance (conductor loss), leakage reactance, 
eddy-current loss, and hysteresis loss. This section will 
expand upon the ideal-transformer concept to produce a 
transformer equivalent circuit that accounts for these 
losses and is a good approximation for real-world trans- 
former performance under any condition. 

Conductor Loss 

As the conductors used for the transformer windings 
have resistance, current flowing in the primary and sec- 
ondary produces an I 2 R power loss that creates heat. The 
loss is minimized by conductors with larger cross sections, 
but if the resistance is too large to be neglected, primary 
resistance, R l7 and secondary resistance, R^, can be placed 
in series with the ideal-transformer windings as shown in 
figure 3.12. 

Leakage Reactance 

For the ideal transformer, all the flux produced by the 
primary must link with the secondary winding. In the real 
world, however, a small percentage of the total flux pro- 
duced fails to link all the secondary turns; this is called 
leakage flux. Leakage flux can be reduced by placing the 
primary and secondary windings very close together, per- 
haps interleaving them. Further reduction comes from 




Figure 3.12.— Ideal transformer with winding resistance in- 
cluded. 



69 



winding the coils tightly on the core and providing a short 
magnetic path between them, thus creating a low- 
reluctance path between the coils. Nevertheless, even with 
the best transformer designs, leakage is significant and 
cannot be neglected. 

Inductance is the ratio of flux linkage to the current 
producing the flux, or 



L = 



Nd<ft 
di 



(3.29) 



where d<£ = magnetic flux, Wb, 

Nd</> = flux linkage of circuit, Wb, 
and di = current producing flux, A. 

For transformers with iron or ferromagnetic cores, current 
and flux do not have a linear relationship, and differen- 
tials must be used. Consider the time-varying primary 
current, i lt in figure 3.13A, where the changing current 
produces the magnetic flux, <^> 1 , and 



Li = 



Nid0i 
dix 



(3.30a) 



The part of <t> 1 that links the secondary is <t> 12 ; that which 
only links the primary (or is lost in terms of magnetic 
coupling) is 0l X , where 



Similarly, although not shown in the figure, 



N 2 d<ft 2 
Lz= di, 



02 = 021 + 0L2> 



L 2 = 



N 2 d<£ 21 N 2 d<^. 1 



di.. 



die 



(3.31a) 
(3.316) 

(3.31c) 



Interestingly, the coefficient of coupling is also related to 
flux by 



k _^12 

01 



021 
02 



(3.32) 



The first term in equations 3.30c and 3.31c is the 
transformer mutual inductance, and the second terms are 
the primary and secondary leakage inductances, L L1 and 
Ll2> respectively, or 



M = 



Nid^ = N 2 d<ft 2 
clii 



di 9 



(3.33a) 



01 = 012 + 0L1 



and 



Li = 



N]d<fr 12 _ Njd^j. 
di x dii 



(3.306) 



(3.30c) 



Lm = 



■Lit 9 — 



di! 



N 2 d^L 2 
di, 



(3.336) 



(3.33c) 




Load| fv L 



M 



R 



L, 




l : a_ L -|_2 R2 



1 



v~ 



v, 



Figure 3.13.— Accounting for transformer leakage flux. 



These equations hold for effective sinusoidal current. In 
steady ac analysis, the leakage inductances become leak- 
age reactances; hence, flux leakage can be represented as 
an inductance or reactance. Figure 3.13S shows the addi- 
tional elements that bring the transformer model closer to 
a practical transformer. 

Core Losses and Exciting Current 

Even with the addition of winding resistance and 
leakage reactance, equation 3.276 for an ideal transformer 
still applies and can be rewritten as 

Ii = al 2 . 

Examination of this expression suggests that whenever I 2 
is zero, I x must be zero. Yet, if an actual transformer 
primary is connected to an ac source and the secondary is 
left unconnected (fig. 3.14), the primary current will exist, 
albeit very small. Even though the secondary is open and 
I 2 is zero, V 2 appears across the secondary winding as a 
sinusoid. This implies that a changing flux in the trans- 
former core must be produced by the current in the 
primary, as no other sources of changing flux are avail- 
able. The portion of primary current that produces the 
changing flux, called magnetizing current, i^, can be 
accounted for by adding an inductor, L e , in parallel with 
the ideal-transformer primary winding. 



70 



The changing flux also induces small currents, eddy 
currents, in the transformer core material. These have an 
almost infinite number of closed paths and encircle prac- 
tically all the flux. Since the transformer core has electri- 
cal resistance, the result is heat in the core and attendant 
power loss. These eddy currents flow at right angles to the 
magnetic field, as illustrated in the core cross section of 
figure 3.15A. The resistance along the eddy-current path 
is approximately proportional to the path length. Obvi- 
ously, if the path length is decreased, the power dissipated 
in the eddy-current loop will drop. Figure 3.15S shows the 
core split lengthwise with a nonconducting layer between 
the two halves. The result is a desirable decrease in power 
loss to about two-thirds of the original power. In practice, 
transformers are laminated from several thin sheets of 
steel. Each sheet is sometimes covered with varnish to act 
as an insulant, but in most cases the oxide layer on each 
steel sheet is sufficient to produce the necessary high- 
resistance layers. This can substantially reduce eddy cur- 
rents but cannot completely eliminate them. 

As shown in figure 3.15C, energy is also dissipated in 
the transformer core each time a hysteresis loop is tra- 
versed. The energy is proportional to the area enclosed in 
the hysteresis loop and is called the hysteresis loss of the 
transformer core. In simple terms, the effect is related to 
the fact that the core retains some magnetism, and a 
coercive force is required to overcome this residual mag- 
netism each time the ac current reverses. The loss is due to 
retentivity or molecular friction. 

Both eddy-current and hysteresis losses are propor- 
tional to frequency and become a major consideration in 
high-frequency transformer applications. However, these 
core losses can be satisfactorily approximated at one 
frequency and one voltage. Good examples are 60-Hz 
power transformers where neither frequency nor voltage 
(actually, magnetic saturation of the core) changes drasti- 
cally in normal operation. To account for these losses, a 
resistance, R e , is again placed in parallel with the ideal- 
transformer primary. The sum of the currents through R e 
and L e is called exciting current, I e , and the total current 
drawn by the source when the transformer is supplying 
power to a load is I x + I e . 

It should be noted that with sinusoidal input voltage 
to the primary, the exciting current is not a sinusoid but 
exhibits many harmonic frequencies because of the 
greatly varying permeability of the transformer core. 
However, for most purposes it may be assumed as a 
sinusoid with the same rms value. 

The equivalent circuit shown in figure 3.16 now 
contains all the components necessary for it to be a useful 
model of a practical transformer. In summary, the impor- 
tant parameters for an equivalent circuit are 

• R x , primary conductor resistance, 

• Ro, secondary conductor resistance, 

• L L1 , primary leakage inductance, 

• L L £, secondary leakage inductance, 

• R e , a resistance accounting for eddy-current and 
hysteresis losses, 

• L e , an inductance accounting for magnetizing cur- 
rent, and 

• An ideal transformer with turns ratio, a = Ng/N^ 




l:a 



L L2 R 2 
-»-'' YYV ^-*- y /\/W-» 



4 VliiPi 

lJ v 'I III 



i 2 =0 



i m = magnetizing current 



Figure 3.14.— Transformer magnetizing current. 




Eddy-current 
/ path 



Eddy-current 
path X 



Flux 



Nonconducting 
layer 

Eddy-current path- 




Energy 

dissipated 

in core 



Magnetizing 
force 



B 



Figure 3.15.— Eddy current 
creating power loss in core. 



and magnetic hysteresis 




N, N 2 
Figure 3.16.— Equivalent circuit of practical transformer. 



Notice that mutual inductance, M, does not appear in the 
model since it is represented by the turns ratio of the ideal 
transformer. 

Power-Transformer Construction 

The two most widely used transformer types are the 
core and the shell. In shell construction, both primary and 
secondary windings are placed on an inner leg of the core. 
The windings are constructed in layers with an insulating 
barrier between them, forming a very low-leakage flux. In 
core construction, the primary and secondary windings 
are located on separate legs, thus providing maximum 
isolation between the coils. Both constructions are 
sketched in figure 3.17. 

A copper or aluminum conductor is employed to 
construct each winding, which can have the form of an 



71 



insulated wire with circular or rectangular cross section, 
or an uninsulated wide metal sheet. The insulated wire is 
continuously wound in layers, with each layer separated 
by a sheet of insulating material. With sheet-metal wind- 
ings, the conductor is wound simultaneously with a con- 
tinuous sheet of insulating material so that each adjacent 
conductor turn is separated by the insulation. The sheet 
metal is the same width as the transformer winding, and 
the insulation sheet is slightly wider. 

Each winding is given a rated capacity, a rated 
current, and a rated voltage. These ratings depend upon 
the number of turns in the winding, the magnetic inter- 
action with other windings, the current-carrying ability of 
the conductor, as well as the ability to dissipate heat 
through the insulation to the environment surrounding 
the winding. It should be obvious that the rated capacity, 
current, and voltage are mathematically related. 



Secondary 



Primary 




Primary 
Secondary 




Core construction 



Core 



Shell construction 



Figure 3.17.— Common power-transformer construction 
techniques. 



Transformer Models 

Since voltage regulation, efficiency, and heating are of 
prime importance in mine power systems, detailed power- 
transformer analysis requires consideration of the com- 
plete equivalent circuit as shown in figure 3.16. However, 
because the exciting current, I e , is normally very small 
compared with load current, I x , a further approximation 
can be made by placing R e and L e at the transformer input 
terminals (fig. 3.18). This modification now allows the 
secondary winding resistance and leakage inductance to 
be transferred to the primary circuit (fig. 3.19) and com- 
bined with the primary elements. For many purposes, the 
exciting current is so small that R e and L e can be removed 
from the model. Figure 3.20 provides this last simplifica- 
tion, where the winding resistance and leakage reactance 
are said to be referred to the primary, and 



R = Ri + ~2 



andl^ = L L1 + — g- 



(3.34) 



The primary is sometimes called the high side if its 
winding has a greater voltage rating (or more turns) than 
the secondary. The secondary is then called the low side. 
The terminology is reversed if the secondary has the 
higher voltage. In steady ac analysis, the inductance 
becomes a reactance, X L , and 




Figure 3.18.— Movement of exciting components to input. 



L_l2 R2 
Rl Uu a 2 a 2 

*-VVV\r^ y ^ v ^^-^- rY ^ v ^ , -V\AAr-^ 



Re 




Figure 3.19.— Transferring secondary components to 
primary. 



R = R x + 



R 



I and X L = w(L L1 



with the primary impedance simply 
Z = R + jX L . 



J L2 
„2 



), (3.35a) 



(3.356) 



If desired, the primary impedance can be moved to the 
secondary of the ideal transformer (thus, referred to the 
secondary) by multiplying both terms by a 2 . 



R k r-a 



Figure 3.20.— Final simplification of practical circuit model. 



72 



EXAMPLE 3.7 

A two-winding transformer has a rated capacity, 
primary-winding voltage, secondary-winding volt- 
age, and frequency of 100 kVA, 2,400 V, 240 V, and 
60 Hz, respectively The primary-winding imped- 
ance is 0.6 + j0.8 A, while the impedance of the 
secondary winding is 0.005 + jO.007 fi. The trans- 
former is being used at the end of a feeder to step 
down voltage to a load. The feeder impedance is 0.05 
+ jO.l Q, and the load is 0.3 + j0.4 Q. Find the 
magnitude of the voltage across the load if the 
voltage at the source end of the feeder is held 
constant at 2,400 V. 

SOLUTION. As core-loss and magnetizing-current 
elements are not given for the transformer, they 
must be assumed to be negligible, with the trans- 
former model being as shown in figure 3.20. The 
turns ratio is 1/10, and the transformer impedance 
is (equation 3.35) 



Z T = R x + jX 1 + —s + j — 2 



a^ 



n „ . no 0.005 .0.007 

- °- 6 + j0 - 8 + am + 3 wmf 

= 1.1 + jl.5 fi. 

The load impedance transferred to the primary is 
Z L 0.3 + j0.4 



(l/10r 



= 30 + j40 Q, 



and the total impedance at the source end of the 
feeder is 



Z eq - Z f + Z T + 2 

= 0.05 + jO.l + 1.1 + jl.5 + 30 + j40 
= 31.15 + J41.6 = 51.97| 53.2° fi. 

The magnitude of current from the source is then 



Ii = 



2,400 



Z eq | 51.97 



= 46.18 A, 



which is also the current through the transformer 
primary. Therefore, the magnitude of voltage across 
the primary is 



V x - I x |% | = 46.18(50) = 2,309 V, 



and that across the secondary and the load is 
V 2 = V x a = ^ = 231 V. 



Determination of Transformer Parameters 

Two tests can provide the necessary elements for the 
transformer model in figure 3.19, where the exciting 
current components are at the primary terminals, and 
secondary parameters are referred to the primary. 

The first test, the open-circuit or excitation test, is 
used to find the exciting-current components. The second- 
ary of the transformer is unconnected. Rated voltage at 
rated frequency is applied to the primary winding, and a 
wattmeter (see chapter 5) is employed to measure the 
power, P e , delivered by the source. An ammeter is used to 
measure rms exciting current, I e (fig. 3.21A). The power 
corresponds to the core loss; in other words, P e is dissi- 
pated by R e . R e can be found by 



V 2 
R e = p- , (3.36) 



where V = rms applied voltage, V, 

P e = measured average power, W, 
and R e = core-loss resistance, B. 

From the rms value of I e , assuming exciting current to be 
a sinusoid, the input admittance of R e and L e in parallel 
can be calculated from 



Realizing that 



Y = — 
e v 



Y = -1 + J- 

6 Re J«V 



(3.37a) 



(3.376) 



the component accounting for magnetizing current, L e , 
can be obtained from 



coL_ 



R„ 



(3.37c) 



where I e = measured rms value of exciting current, A, 

w = 27rf = applied frequency of source (must be 

rated frequency of transformer for exact 

results), 
and L e = magnetizing inductance, H. 

Therefore, both R e and L e can be determined from the 
open-circuit excitation test. 

The second test is projected at winding resistance and 
leakage resistance, with both primary and secondary 





A Open circuit B Short circuit 

Figure 3.21.— Transformer parameter test series. 



73 



values combined. This is termed the short-circuit or im- 
pedance test. Here, the secondary terminals are short- 
circuited and a source is connected to the primary. Voltage 
at rated frequency is applied to the transformer but at 
reduced amplitude, so that it produces only rated current 
in the primary winding and, thus, rated current in the 
secondary. Current, 1^, and input average power, P^., are 
again measured. 

The applied voltage for the test is typically much 
smaller than rated voltage. Yet the short-circuit (actually, 
rated) current is much greater than the exciting current, 
so I e and the associated components can be neglected. As 
given by figure 3.215, the equivalent circuit under these 
conditions can be simplified to a simple series RL combi- 
nation. The ideal transformer is not needed because the 
zero load impedance (short circuit), when transferred to 
the primary, is still zero. Winding resistance and leakage 
inductance can thus be found from 



often applied to transformer secondary-voltage variations 
and is defined as 



V.R. = 



Vnl V fl (1Q0%)) 



(3.41) 



where V FL = transformer output voltage at full rated 
secondary current and rated primary volt- 
age, V, 

and Vnl = transformer output voltage with no second- 
ary load but rated primary voltage applied, 
V. 

V FL and Vnl are also called the full-load and no-load 
voltages, respectively. It should be clear that voltage 
regulation is a function of transformer losses, impedance, 
and efficiency. The concept is extremely important in mine 
power systems as it often limits how far a mine can be 
safely expanded from one power source. 



R = 



(3.38) 



and 



wL l = (^2" 



,V 2 = 
I 2 



-R 2 ) 1 



(3.39) 



where Pg,. = measured average power, W, 

1^, = measured rms short-circuit current, A, 
Vs,. = applied rms short-circuit potential, V, 
co = 2irf = rated frequency, rad/s, 
R = primary and secondary winding resistance, 
12, 
and L L = primary and secondary leakage inductance, 
H. 

It is important to note that these values are valid only 
for the frequency under which the tests are made. Further, 
it is neither possible nor necessary to break the resulting 
components into primary and secondary elements. 



Transformer Efficiency and Regulation 

The transformer is designed to be a highly efficient 
device. However, the output power of a transformer is 
always less than its input power because of winding 
conductor losses and core losses. The term efficiency is 
used to measure the ability of a transformer to transfer 
energy from the primary circuit to the secondary circuit. 
The efficiency is defined as the average-power ratio: 



V = 



(3.40a) 



EXAMPLE 3.8 

For the circuit shown in figure 3.22, find the 
complex power consumed by the transformer load, 
Z L . If the figure represents the full-load condition, 
what is the voltage regulation at the transformer 
secondary? The transformer is considered ideal. 

SOLUTION. The impedance seen by the 5,000-V 
source is 



Z eq = 1 + jl + 



= 1 + jl + 



(0.1 + J0.1) 



(1/10) 2 
= 11 + jll = 15.56|45^ Q. 

Using the source voltage as the reference phasor, the 
current delivered from the source is 



_ 5,000|0°, _ 3212| _ 45 o A 
11 _ 15.56145° " 321.2|_45_ A, 



and the transformer secondary current is 



l 2 = h = 10(321.31 -45° ) = 321.31 -45° A. 
a 



P ; „ - losses 



or 



V = 



+ losses 



(3.406) 



The ratio is always less than 1 but normally in the range 
n = 0.95 to 0.98. Efficiency decreases when the device is 
operated above or below its voltampere capacity. 

Voltage regulation is a characteristic of power systems 
that describes the voltage fluctuations resulting from 
varying load or current conditions. Voltage regulation is 



in jin 10:1 J2, 



5,000 V, 
60 Hz 




z L =o.i+jom 



Figure 3.22.— Circuit for example 3.8. 



74 



The voltage across the transformer secondary is 

V 2 = I 2 Z L = (3,213| -45° X0.141|45°) 
= 454.4|0^ V, 

and the complex power delivered to Z L is 

§l = V 2 I 2 * = (454.4|0°X3,213|45°); 

therefore, Sl = 1.46 1 45° MVA. 

If the above situation represents the full-load condi- 
tion, then 



V FL = V 2 = 454.4 V. 



Under no-load conditions, the load impedance be- 
comes such a high impedance that the transformer 
secondary current approaches zero. With no second- 
ary current, current to the primary of an ideal 
transformer is also zero. Therefore, the voltage 
across the primary is equal to the source voltage or 

Vj. = V 8 = 5,000 V. 

The secondary voltage becomes 

5,000 
V 2 = aV 1 = -^- = 500 V. 

Consequently, Vnl = V 2 = 500 V, 
and from equation 3.41, 

500 - 454 



V.R. = 



454 



(100) = 10%. 



— *■ \-a — ' 



AUTOTRANSFORMERS 

All the transformers discussed so far have been two- 
winding transformers and have provided electrical isola- 
tion between the primary and secondary windings. An- 
other type of transformer, the autotransformer, uses a 
single winding and does not provide electrical isolation. It 
is constructed from a continuous winding with a tap 
connected at a specific point. The autotransformer is 
compared with an ideal two-winding transformer in figure 

3.23. The advantages and disadvantages of each type of 
transformer can be illustrated with reference to figure 

3.24, where a normal two-winding transformer is shown 
on the left and is connected to operate as an autotrans- 
former. 

The two-winding transformer has the following spec- 
ifications: 

N x , Vj, Ij = primary turns, rated rms voltage, and rated 

rms current, 
N 2 , V 2 , I 2 = secondary turns, rated rms voltage, and 
rated rms current, 

and the maximum apparent power that can be delivered to 
a secondary load is 

S„,„ = VJ 2 . 



t ' ' t 



N, N 2 
A 



B 



Figure 3.23.— Comparison of two-winding transformer {A) 
and autotransformer (8). 







v, 



N, 



A 



B 



Figure 3.24. — Two-winding transformer as an 
autotransformer. 



To help visualize the autotransformer action, figure 
3.24A is redrawn in figure 3.24ZJ with both windings 
placed on the same side of the core symbol. For either 
figure, the output voltage, V 2 ', is now 

V 2 ' = v, + v 2 . 

Transformer rated output current, I 2 , is still related to 
rated primary current, I 1; by 

Ii = al 2 , 

but input current to the autotransformer is now 

Ii' = Ii + I 2 - 

The maximum power that can be transferred to a load at 
rated output current, I 2 , is now 

s out = V 2 'I 2 + V X I 2 + V 2 I 2 . 

This expression indicates that the transformer is now 
able to deliver an increase of V X I 2 voltamperes over the 
two-winding connection, yet the transformer windings are 
still within rated currents and voltages. The reason for the 
increase is that some input current is transformed by the 
transformer while the rest is conducted directly to the 
load. This is the main advantage of the autotransformer 
over two-winding arrangements. Because primary current 
is now only a portion of load current, conductor losses in 
autotransformers are particularly small, and voltage reg- 
ulation under varying load conditions is usually good. 



MULTIVOLTAGE TRANSFORMERS 

The transformers considered so far have had only one 
secondary, but in practice many have two or more second- 
ary windings. The transformer with two secondary wind- 
ings in figure 3.25A is able to serve loads with different 



75 



voltage requirements from one source. In such devices the 
magnetic interaction increases substantially over the two- 
winding variety because mutual inductance exists be- 
tween all winding combinations. Taking this into account, 
the preceding theory can be expanded to model an equiv- 
alent circuit. 

Another method for one transformer to serve several 
voltage applications is to have winding taps on the pri- 
mary (fig. 3.25B), the secondary (fig. 3.250, or both. When 
used on the input winding, a higher tap can be selected to 
account for voltage drops in the circuit that delivers power 
to the transformer, thus maintaining a desired output 
voltage. This is a common practice in mining. A special 
but very widely used application for secondary taps is in 
utility distribution transformers supplying 240- and 120-V 
ac service. Here, the winding is center-tapped with equal 
turns on either side. The voltage magnitude from either 
line to the tap is 120 V, and across the total winding, 240 
V is available. 



CURRENT AND POTENTIAL TRANSFORMERS 

The prime use of transformers in mine power systems 
should now be apparent: to supply power at different 
voltage levels to system portions and equipment. Trans- 
formers are also used extensively to power control circuits, 
mainly to provide power for circuit breakers and associ- 
ated circuitry; to power protection devices, usually relays 
to trip circuit breakers; and for instrumentation. Trans- 
formers employed for these applications are often given 
specific names: potential transformers (PT's) and current 
transformers (CT's). PT's are merely high-quality two- 
winding transformers with or without taps. The name is 
modified because they are used to sense voltage. 

The current supplied to relays, instruments, and sim- 
ilar equipment is normally provided by CT's. Some CT's 
are like the two-winding devices that have just received so 
much attention. These have a primary with just a few 
turns of high-current-capacity conductor and a secondary 
with numerous turns, as illustrated in figure 3.26A. The 
turns ratio C^/N^) is normally adjusted so that the 
secondary supplies 5 A when full-load current flows in the 
primary. The primary is placed in series with the circuit 
that is to be measured, and therefore, CT's can be consid- 
ered as sensing current. 

Two-winding CT's for high-voltage or high-current 
circuits, such as those usually found in mine power sys- 
tems, are very expensive, and as a result bushing-type or 
donut CT's are more often used. In figure 3.26B, the 
conductor to be measured passes through a large-diameter 
ring-shaped laminated iron core and acts as the trans- 
former primary. The secondary winding, which consists of 
several turns about the core, supplies current as before. 
The leakage reactance of this type of CT is obviously high 
and, coupled with other parameters, results in a low 
accuracy for current measurements. A schematic illustrat- 
ing hypothetical placements of a PT and a CT in a simple 
circuit is provided in figure 3.27. 

PT's and CT's are important components in instru- 
mentation and protective circuitry for mine power sys- 
tems. Their application for instrumentation is presented 



Extra taps 





Figure 3.25.— Examples of transformers for multi voltage ap- 
plications. 



Primary : a few turns 
of high-current 
conductor 



Power 
conductor 



In series 
with circuit 



To instruments 
and relays 



N, IM 2 

Secondary: several 
turns of 
conductor 




To instruments 
and relays 



2- winding transformer Bushing or donut transformer 

Figure 3.26.— Two types of CT's. 



I Source 



Load 






CT 

n 



pr * — Senses I I ' v -- Senses 

Pl voltage * " current 

To instruments 
or relays 



To instruments 
or relays 

Figure 3.27.— Examples of CT and PT placement in circuit. 



in chapter 5, while chapters 9 and 10 cover their use in 
protective relaying. 

The purpose of the foregoing two chapters was to cover 
many of the basic theoretical aspects behind mine electrical 
systems. The content was directed towards dc and single- 
phase ac, and spanned fundamental electrical phenomena, 
the experimental laws and parameters, dc and ac circuit 
analysis, and finally, power transformers. Comprehension of 
these laws, parameters, and concepts is essential for the 
understanding of subsequent chapters. This will be very 
apparent in the next chapter, which introduces power-system 
concepts and three-phase circuit basics. 



76 



CHAPTER 4.— POWER-SYSTEM CONCEPTS 



Power systems can be simply described as systems 
that transmit power from a source to the loads. For the 
mine, the source is often the secondary of a substation 
transformer and the loads are motors on mining machin- 
ery and ancillary equipment. The transmission of power is 
commonly performed by three-phase systems, which are by 
nature more complex than the dc and single-phase ac 
circuits introduced in the previous two chapters. The 
following sections are primarily concerned with three- 
phase power systems plus the basic tools and special 
mathematics needed to study them. Several references are 
provided at the end of the chapter. As most information is 
considered common electrical engineering knowledge, spe- 
cific references are seldom cited but can be found in the 
bibliography. 



BASIC POWER CIRCUIT 



THREE-PHASE CIRCUITS 

The term single phase has been applied to ac 
systems where power is delivered from a single sinusoidal 
source. When power is transmitted to a load by applying 
two or more sinusoidal sources with fixed phase differ- 
ences, the power system is called polyphase. The most 
popular system that delivers large quantities of power, 
including both single phase and polyphase, is the three- 
phase system. 

The analysis of three-phase circuits can be extremely 
complicated. Special techniques have been developed to 
assist in general problem solutions, but even so, the work 
can be cumbersome. However, three-phase systems are 
purposely designed to be balanced, and if actual differ- 
ences existing among phases can be neglected, the analy- 
sis of three-phase circuits can be almost as simple as 
analysis of single-phase circuits. 



Many power systems or system segments can be reduced 
to the simple series circuit shown in figure 4.1. This familiar 
single-phase ac circuit consists of a source or supply voltage, 
an impedance, and a load or receiver voltage. Such a repre- 
sentation is often called the Thevenin's equivalent of the 
power system. Finding the series circuit may involve many 
simplifying assumptions or procedures, some of which are 
yet to be covered, but the result has numerous applications 
for analyzing the behavior of electrical power systems. 

One specific example is analysis of voltage regulation. 
Here, the source voltage is kept constant, and variations of 
the load voltage are observed with a range of load-current 
conditions that cause a change in voltage drop across the 
impedance. Applying this example to an underground coal 
mine, the source could be the secondary of a power-center 
transformer, the impedance could be that of the trailing 
cable, and the load might be the motors of a continuous 
miner. On a larger scale, a substation output voltage, a 
feeder cable, and power-center primary voltages could 
constitute a desired Thevenin's equivalent for analysis. 
Both these situations are illustrated in figure 4.2. As the 
chapter unfolds, more applications will become apparent. 

Actual analysis of the basic power circuit (fig. 4.1) can 
use any applicable technique already given in chapters 2 
and 3. For instance, employing the impedance domain and 
Kirchhoff s voltage law yields 



BALANCED THREE-PHASE CIRCUITS 

Balanced three-phase power consists of three gener- 
ated voltages, each of equal magnitude and frequency but 
separated by 120°. When these voltages are applied to a 
system of balanced impedances, balanced currents result. 
In other words, a balanced three-phase power system can 
be divided into three portions. Any voltage or current in 
one portion has a counterpart in another portion, which is 
identical but 120° out of phase. 



Source or 
supply 




Load or 
receiver 



Figure 4.1. —Basic power circuit. 





V 8 - IZ - V L = 


(4.1a) 


or 


V 8 = V L + IZ 


(4.16) 


and 


v L = v 8 - iz. 


(4.1c) 




Continuous 
miner 



Trailing cable 



SOURCE 



^S^ 



IMPEDANCE 



LOAD 



Any variable or constant in these equations can be a 
complex expression. Nevertheless, the equations describe 
the performance of the power system that the circuit 
represents, that is, the source voltage for a specific load 
current and load voltage, and so forth. When three-phase 
systems are involved, the solution or even the finding of 
the equivalent circuit must also utilize the additional 
methods that follow. 



Substation 



Feeder cable 




SOURCE IMPEDANCE LOAD 

Figure 4.2.— Applications of basic power circuit. 



77 



lb illustrate this voltage generation, consider the 
elementary three-phase generator illustrated in figure 
4.3A. The armature consists of three single stationary 
conductors displaced by 120°, and a magnetic field struc- 
ture rotates counterclockwise within. As the rotating 
magnetic flux cuts each winding, a voltage is induced. 
These voltages are out of phase with one another, as shown 
in figure 4.3B. A composite of these instantaneous volt- 
ages is provided in figure 4.3C to exemplify the phase 
relationships, which also can be clarified with phasors (fig. 
4.3D). It can be noted with either representation that the 
voltage in winding aa' reaches a maximum first, followed 
by bb', and then cc'. This defines the positive sequence, 
abcab. . ., that is evident from the counterclockwise ro- 
tating phasors of figure 4.3D. If the phasors are allowed to 
rotate in the opposite direction (clockwise), the sequence 
termed negative (cbacba . . .). 

An outstanding advantage of balanced three-phase 
systems is that they provide a more uniform flow of energy 
than single-phase or even two-phase systems. The 120° 
timing means that the individual power waves in each 
phase never reach zero at the same time, and more 
important, the total instantaneous power from all three 
phases remains constant. For three-phase motors, this 
translates to convenient starting, constant torque, and low 
vibration. It would seem logical that if three phases 
provide a substantial increase in operation efficiency, more 
equally spaced phases would result in even greater im- 
provement. However, three-phase systems are generally 
more economical than other polyphase systems because 
the complications caused by additional phases offset the 
slight efficiency increase. 

A source supplying these three-phase voltages is nor- 
mally connected in either delta or wye. As shown in figure 
4.4, either configuration can, in practice, be closely ap- 
proximated by ideal voltage sources or in some cases by 
ideal voltage sources in series with small internal imped- 
ances. Three-phase sources always have three terminals, 
which are called line terminals, but may also have a fourth 
terminal, the neutral connection. These terminals produce 
three separate potentials between any two line terminals 
that are called line-to-line voltages. Also generated are 
three separate voltages between each line terminal and 
the neutral, be it a direct connection as in figure 4.4 or 
some imaginary neutral point. These are termed line- 
to-neutral potentials. 




Single conductor 
armature winding 

A Generator 



V QQ ' V bb < V cc < 




B Individual waveforms 





I- — 4 — A 
120° 120° 

C Combined waveforms D Voltage phasors 

Figure 4.3.— Elementary three-phase generation. 



Line terminal 







^Line conductor 








a 


}b 

Neutral 
conductor 


Line "\ "J< Line-to- 
to-hneXfc) neuitm 

voltage V^rt. } voltage 



Figure 4.4.— Three-phase voltage sources. 



Three-Phase System Voltages 



Line-to-line voltage can be considered as a condition 
existing between two phases, while line-to-neutral is a 
condition for one phase only. Obviously, interrelationships 
must exist between these two voltage notations, as well as 
among the voltages of one notation. The wye-connected 
source of figure 4.5A can be employed to demonstrate the 
correspondence. 

If the line-to-neutral voltages, V an , V^, and V cn , are 
positive sequences and _the phasor of V^ is taken as 
reference, then V bn and V^, are related to V an by 

V bn = V.J -120° , V^ = V.J -240« , (4.2a) 





A 


B 


C 


3- phase 


Line-to- line and line- 


Graphical 


wye source 


to-neutral voltage 
phasors 


construction 



or 



V. = VhJ + 120° = V.J -120* 



Figure 4.5.— Wye-connected source demonstrating line-to- 
(4.26) line and line-to-neutral voltages. 



78 



Equations 4.2a and 6 relate that if a specific phasor 
representing one phase voltage is rotated 120°, it is 
identical to the phasor for another phase. By Kirchhoff s 
voltage law, the line-to-line voltage is equal to the sum of 
the two line-to-neutral voltages; for instance, between 
phases a and b, 





v ab = v an + v nb , 


(4.3a) 


but 


v nb = -v bn 




and 


Vk„ = V.J-120-; 




hence, 


V ab = V an - V an |-120° 


(4.36) 


or 


V ah = V3V a J+30°. 


(4.3c) 



loads are assumed to have no impedance, although obvi- 
ously, in the real world, they must have impedance. 

Figure 4.6A shows the first arrangement to be consid- 
ered, the four- wire wye to wye. The source here could be 
either a generator or the secondaries of an ideal three- 
phase transformer, and the load, Z a , Z b , Z c , could be a 
motor. These conductors are connected between the source 
line terminals and the load; the fourth conductor, the 
neutral return (or just simply, the neutral), connects the 
neutral of the source to the common junction of the three 
load impedances. 

For perfect conditions, the generation is balanced, 
distribution impedances per phase (again assumed zero 
here) are equal, and the load impedance in each phase 
circuit is identical. Hence, the magnitude of the line 
currents, I a , I b , and I c , must also be equal. By Kirchhoff s 



Equation 4.3c is truly significant because it states the 
relationship between line-to-line and line-to-neutral volt- 
ages for balanced three-phase systems. In particular, the 
following can be extracted: 



and 



|V ab | = |V bc | = |V ac | 
= V3 |V an | = V3 |V bn | = V3 |V C 



IV. 



= IVkJ = IV, 



bn 



1 

V3 



1 
V3 



v ab | = j* |V bc I = -h IV, 



1 

V3 



(4.4a) 



(4.46) 




Neutral foj) VQn 
point v^rn ^\- 



Vbn 



^ab 



lb. 



Ir 



-Neutral return 




I a + Ib + I c 



A 4-wire wye-to-wye 



It is important to note that, in addition to the foregoing 
identities, for a balanced three-phase system, 



v an + v bn + v cn = 



and 



V 



ah 



+ v bc + V ca = 0. 



(4.5a) 
(4.56) 



A phasor diagram illustrating all line-to-line and line- 
to-neutral voltages of these systems is given in figure 4.5S. 
Here the correspondence by equation 4.3c is apparent. 

The reasoning used for voltages can be applied to 
currents, and this will be handled shortly. 



Load Connections 

As with sources, balanced three-phase loads can be 
connected delta or wye. However, the interest in three- 
phase circuits comes from how delta or wye sources supply 
power to delta or wye loads. The usual combinations or 
systems are 

• Four-wire, wye to wye; 

• Three-wire, wye to wye; 

• Three- wire, wye to delta; 

• Delta to delta; and 

• Four-wire, wye to delta. 

By analyzing each combination, certain advantages and 
disadvantages can be seen, and some important points 
about balanced three-phase systems can be gained. For 
purposes of discussion, the lines connecting sources to 




;+ _Rb jb_^ 




B 3- wire wye -to- wye 




^ n V bn fv ab T^ 



C Wye- to -delta 




b Kb ib, 





D Delta- to -delta 
Figure 4.6.— Balanced three-phase load connections. 



79 



current law and the 120° displacement of the three line 
currents, the neutral-return current must be 



I a + I b + I c = 0, 



(4.6) 



which means the neutral conductor actually carries no 
current under this ideal situation. Furthermore, there will 
be no voltage drop across the neutral, no matter what the 
neutral impedance is. In other words, the potential at the 
neutral of the source equals that of the load. 

If the neutral carries no current under balanced 
conditions, what purpose does it really serve and can it be 
removed? Consider figure 4.6B, a three-wire wye-to-wye 
system, which does not employ the neutral conductor. 
Although this system is used in some applications, prob- 
lems can arise, and the role of the neutral conductor is to 
minimize these problems. 

In the real world, no balanced three-phase system can 
be perfect, and the sources, the distribution impedances, 
and the loads can easily become unbalanced, that is, 
unequal from phase to phase. The result is unbalanced 
currents and voltages. For example, without the neutral 
conductor, the neutral of the source will not equal the load 
neutral, and the resulting load unbalance will produce 
unequal voltages across the loads, no matter how balanced 
the source. Under this condition, a mining machine motor 
is likely to deteriorate and the result will be maintenance 
problems. In addition, safety problems can abound as a 
result of the unequal neutral potentials alone. Chapter 7 
will investigate many of these problems in detail. 

It is apparent that the neutral conductor does serve a 
vital role in actual three-phase power systems. Its size and 
current-carrying ability do not need to match those of the 
phase conductors in order to provide the necessary func- 
tion. In a properly operating power system, normal condi- 
tions do cause some neutral current, but this is usually 
very small compared with the phase current. Hence, 
neutral conductors could be small if they were based only 
on the size of the neutral current, but in mining applica- 
tions, this is not the only criterion. Possible system mal- 
functions must also be taken into account, and these will 
be discussed in a later section on unbalanced three-phase 
circuits. 

A three-phase load is more likely to be delta connected 
than wye connected. The three-wire wye-to-delta system, 
shown in figure 4.6C, is an example of this arrangement. 
The prime advantage is that under unbalanced load 
conditions, the source will deliver power proportionately 
to each load. Hence the delta-connected loads need not be 
precisely balanced. Flexibility is increased because phase- 
to-phase loads may be added or removed without signifi- 
cantly upsetting system operation. With wye-connected 
loads such changes are difficult or nearly impossible to 
make. 

A delta-connected source is shown in figure 4.6D. 
Although this arrangement can be found, it has two major 
disadvantages. First, a slight unbalance in the source can 
create large circulating currents around a delta loop (for 
example, source V ab and load Z ab ). This extra current can 
reduce the available current capacity of the source and 
also increase power losses in the system. Second, it is more 
difficult for safety purposes to maintain metallic equip- 
ment frames at the neutral potential of the source. The 
logical and most economical point to employ as a ground is 
the neutral of the wye-connected source. This system is 



known as a four-wire wye-to-delta system and is illus- 
trated in figure 4.7. It is presently the most popular 
three-phase power connection arrangement in mining. 
The neutral conductor here is more often termed a ground- 
ing conductor. A neutral point can also be derived from a 
delta source using a zig-zag or grounding transformer (see 
chapter 7). 

Line and Phase Currents 

Currents in a specific phase conductor or in one leg of 
a wye-connected source or load are termed line currents. 
As with line-to-neutral voltages, they can be considered as 
a condition of one phase only. Currents flowing between 
two phases are called phase currents (or line-to-line cur- 
rents) and correspond to line-to-line voltages. An obvious 
example of phase current is that flowing through one leg of 
a delta-connected load. As might be assumed, for the 
balanced three-phase system, the magnitudes of the three 
phase currents through the legs of the delta are equal. 
Figure 4.8A shows a schematic of a balanced delta load 
with three line currents I a , I b , I c and three phase currents 
I ab , I^, I ca . It can be utilized to demonstrate the relation- 
ship between line and phase currents. Considering only 
phase a and using Kirchhoff s current law, 



1q — -IqV. — 1, 



ab -'■ca* 



(4.7a) 



From the same reasoning that related line-to-neutral to 
line-to-line voltages, 



I a = V3 I ab | -30° , 



(4.76) 




Neutral ^h v ™ - 
point - 

w b 



fc Vbn tv ob T^ 



v cn 




^ Neutral return or "grounding conductor" 



I Q +I b +I c =0 
Figure 4.7.— Four-wire wye-to-delta system. 



Neutral 

or ground 

point 





Figure 4.8.— Balanced delta load illustrating phase and line 
currents. 



80 



which means that in the balanced case the magnitude of 
line current is larger than phase current by a factor of 
V3. The phasors are displaced by 30°. The symmetry of 
phase and line currents is shown in figure 4.8B. 



Equivalent Delta and Wye Loads 

There are many instances where it is desirable to 
replace a balanced delta-connected load with a wye, or vice 
versa. The groundwork to perform this change has already 
been established in chapter 3. From equation 2.48, 



z ab z c 



Z an z 



(4.8) 



J ab 



J bc 



and so on for Z,^ and Z cn in terms of the delta impedances. 
Equation 4.8 provides equivalence of delta and wye for all 
situations, including unbalanced loads. However, for bal- 
anced conditions, the expression reduces to simply 



-7 Z ab 

^an~ 3 



or 



z ab = 



3Z C 



(4.9a) 



(4.96) 



This states that each branch of a balanced delta has three 
times the impedance of a balanced wye. 

Now that voltages, current, and equivalent load im- 
pedances of balanced three-phase systems have been cov- 
ered, these values can be compared for delta and wye 
loads. If the load is wye connected, the line current and 
load current per phase are the same, but the voltage across 
each load impedance is line-to-neutral, 1/V3 that of line- 
to-line. When the load is delta connected, the voltage 
across one load impedance is line-to-line, while the line 
current is larger than the phase current through each load 
impedance by a factor of VI. These concepts are illustrated 
in figure 4.9 for equivalent delta and wye loads. It is 
significant to note that the three line-to-line voltages and 
three line currents for either connection are identical. 



Three-Phase Power 

Because the voltage and current are the same in each 
impedance of a balanced delta or wye load, the average 
power consumed by one impedance is one-third of the total 
power to the load. In a delta load as in figure 4.9A, current 
and voltage are phase and line-to-line, respectively, and 



B lr 





If z ab = Z cn - Z bc , V ob - 73 V an , 
Z an = Z ob /3 and 1 a = & J ab 



If Z an =Z bn = Z cn ,V an = V Qb //3, 
z ab = 3Z an and I ab =I a //3 



Figure 4.9.— Comparison of equivalent delta (A) and wye (B) 
loads. 



the angle between them is the angle of impedance. Thus, 
considering phases a and b, the average power consumed 
by one element is 



Ppa = V ab I ab cos0, 



and total power is 



or in general, 



P T = 3 V ab I ab cos0, 



P T = 3 V LL IpCos0, 



(4.10) 



(4.11a) 



(4.116) 



where P pA = average power consumed by each element of 
a delta load, W, 
P T = total power consumed by delta load, W, 
V LL = line-to-line (or system) voltage rms magni- 
tude, V, 
Ip = magnitude of phase rms current through 
load, A, 
and cos0 = power factor of load. 

When the load is wye connected, line current is 
through each load, while the voltage is line-to-neutral. 
Hence, taking phase a (fig. 4.95), the average power to one 
element is 



Ppv = V an I a COS0, 



(4.12) 



and total power is 

P T = 3 V an I a cos0, (4.13a) 

or P T = 3 V Ln I L cos0, (4.136) 

where P py = average power consumed by each element of 
wye load, W, 
P T = total power consumed by wye load, W, 
V Ln = magnitude of line-to-neutral rms voltage, V, 
I L = line rms current magnitude, A, 
and cos0 = power factor of load. 

It is important to note that the power-factor angle, 0, is 
referenced to the sinusoidal voltage across one load and 
the current through that load. 

The standard measurement values for three-phase 
circuits are line-to-line voltage and line current, which are 
often the known quantities. Since for balanced systems, 



and 



V LL = V3 V L 
I L = V3 Ip, 



both equations 4.116 and 4.136 are also identical to 

P T = V3 V LL I L Goad pf), (4.14) 

where the power-factor angle is that of a load impedance or 
an equivalent impedance. It is important to realize that 
this angle has nothing to do with the angle between V LL 
and I L , for example, V ab and I a . Of the three three-phase 
average-power formulas, equation 4.14 is by far the most 
used. 



81 



Following the single-phase presentation of chapter 3, 
a balanced three-phase load has reactive power, Q T , and 
apparent power, St, in addition to average power, P T . The 
following expressions apply: 

Q T = V§ VLLl L sin0 (4.15a) 

or Q T = 3 V Ln I L sin0 (4.156) 

or Q T = 3 VLLlpsinfl (4.15c) 

and &r = V3 V LL I L = 3 V LL I P = 3 V Ln I L . (4.16) 

Complex power, St, is therefore 

Sj, = P T + jQ T = StI», (4.17a) 

or for phase a, 

St = 3 V an I*, (4.176) 

or for phases a and b, 

Sr = 3 V ab I* b . (4.17c) 

It should be evident that in balanced three-phase systems, 
complex power, Sp, does not equal \3 V ab I*. 

Basically, all power concepts presented in chapter 3 
for single-phase ac also apply to balanced three-phase 
power. Poor power factor is worthy of critical attention 
because it affects the entire system operation by limiting 
the available power from transformers, hindering voltage 
regulation, and limiting the current-carrying ability of 
conductors and cables. Simply, the result is poor system 
operation and economy. Thus, power factor can be equated 
to an indicator of system efficiency. 



EXAMPLE 4.1 

An underground coal mining section contains 
the following three-phase equipment connected to 
the secondary of a transformer in the section power 
center: 

• Continuous miner: 300 kW at 0.6 lagging pf, 

• Two shuttle cars: each 60 kW at 0.8 lagging pf, 

• Roof bolter: 50 kW at 0.8 lagging pf, 

• Feeder-breaker: 100 kW at 0.6 lagging pf. 

Find the capacity of the power-center transformer 
necessary to operate these machines. 

SOLUTION. The complex power to several loads is 
the sum of the complex power consumed by each 
individual load. Thus, the complex power for each 
load must be found first. For the continuous miner, 

Pi = Si cos0 a 



or S, = -\ = f^ = 500 kVA, 

COS0 1 0.6 



and Q x = S x sin0 x 

or Q x = 500 (0.8) = 400 kvar. 

Accordingly, for both shuttle cars, 

P 2 = 120 kW, S 2 = 150 kVA, Q 2 = 90 kvar. 

For the roof bolter, 

P 3 = 50 kW, S 3 = 62.5 kVA, Q 3 = 37.5 kvar. 

For the feeder-breaker, 

P 4 = 100 kW, S 4 = 166.7 kVA, Q 4 = 133.3 kvar. 

The total average and reactive power are then 
respectively 

P T = P x + P 2 + P 3 + P 4 

= 300 + 120 + 50 + 100 = 570 kW, 

Q T = Qi + Q 2 + Q 3 + Q 4 

= 400 + 90 + 37.5 + 133.3 = 660.8 kvar. 

The total complex is then simply 

St = P T + jQ T = 570 + J661 kVA 

or Sr = 873|49.22 kVA. 

The required capacity of the transformer is equal to 
the apparent power of the load. Therefore, 

transformer capacity = Sr = 873 kVA. 

It should be noted that, as in chapter 3, the solution 
to the problem cannot be assumed to be the simple 
summation of the apparent powers for all the loads. 
The only case where this is possible is where all 
loads are operating with the same power factor. 



EXAMPLE 4.2 

For the combined consumptions of example 4.1, 
find the necessary total capacitance in kilovoltam- 
peres reactive to improve the overall power factor to 
0.8 lagging. The capacitance will be connected 
across the transformer secondary. 

SOLUTION. The combined complex power for the 
preceding problem is 

Sr = 570 + J661 kVA. 

When pure capacitance is added, average power will 
remain constant, but reactive and apparent power 
will decrease. Therefore, 

cos0 new = 0.8 lagging, 



82 



COS0„ 



570 
= Q-g = 712.5 kVA, 



Hnew — "new Sln "new> 

= (712.5X0.6) = 427.5 kvar. 



The difference between this new or improved reac- 
tive power and that without the capacitance is the 
reactive power less the capacitance, or 

Qc = (Qt " Qnew) = (661 - 427.5) 
= 233.5 kvar. 

It can be noted that this example is much like 
example 3.4. The concept of power-factor improve- 
ment has been repeated here to show the similarity 
of most power problems, be they single phase or 
three phase. 



THREE-PHASE TRANSFORMERS 

Considerable background information about trans- 
formers was presented in chapter 3, and most of that 
theory is also applicable to three-phase transformers. The 
prime purpose is the same as with single-phase systems, to 
provide the different voltages required for distribution and 
equipment operation. The transformer can be constructed 
as either a single three-phase unit or a bank of three 
single-phase units. The only difference between the two is 
that the three-phase unit has all windings placed on a 
common core. 

The connections can best be described by considering 
a bank of three single-phase two-winding transformers. 
Every coil is insulated from the rest, and there are three 
primary and three secondary windings, all of which can be 
interconnected independently. The primary and secondary 
windings can be delta or wye connected while complete 
electrical separation is retained between all the primary 
and secondary windings. The possible connections are wye 
to wye, delta to delta, delta to wye, or wye to delta. Figure 
4.10 illustrates the physical connections of each combina- 
tion, and figure 4.11 shows the corresponding symbols 
used in the three-phase circuit diagram. Any one of these 
combinations can be found in or about mine installations, 
but mine power transformers are typically delta to wye. 

Delta-to-wye connections are popular in mine power 
systems because of the load advantages of the delta 
connection of the primary, which is in essence the load for 
the incoming power. The neutral of the wye-connected 
secondary provides a good grounding point for the outgo- 
ing system from the transformer and does not shift poten- 
tial under unbalanced load conditions. The delta-wye 
winding combination does not generate third-harmonic 
(180 Hz for 60-Hz systems) voltages and currents that 
hamper delta-delta and wye-wye connections. 

The second most popular transformer configuration in 
mines is the delta to delta. Although system s requiring a 
grounding neutral point create some difficulty for the 
delta secondaries, the delta-delta connection has one sub- 
stantial advantage. If one of the single-phase transformers 
fails, operation can be continued by removing the defective 
unit and operating the two remaining transformers as 





Wye- to- wye 



Delta -to -delta 




? a' 




) b 



Delta -to -wye 



Wye -to -delta 



Figure 4.10.— Three single-phase transformers connected 
for three-phase operation. 



Phase 





Wye -to -wye 



Delta- to- delta 




Neutral 




Delta- to- wye 



Wye -to -delta 



Figure 4.11.— Three-phase diagrams for the transformers of 
figure 4.10. 



open delta. This open-delta or V connection can be illus- 
trated by the two single-phase transformers shown in 
figure 4.12. Although it is an unsymmetrical connection, 
it does provide a symmetrical three-phase power input and 
output. However, using the two transformers in this man- 
ner reduces capacity to 57.7% of the three-transformer 
kilovoltampere rating. Nevertheless, it is an effective 
emergency measure. The open-delta configuration is some- 
times used as a temporary circuit; for example, when the 
completion of delta is postponed until load conditions 
warrant a third unit. 



83 




Figure 4.1 2.— Open-delta three-phase transformer operation. 



Calculations with delta-to-delta and wye-to-wye trans- 
formers are straightforward and easy to comprehend. With 
delta to delta, primary line-to-line voltages and phase 
currents are transformed to secondary line-to-line and 
phase values, while for wye-to-wye transformers, line- 
to-neutral voltages and line currents transfer directly. 
Delta-to-wye and wye-to-delta combinations are different. 
With a delta-to-wye configuration, primary line-to-line 
voltages become secondary line-to-neutral, and primary 
phase currents transform to secondary line currents. 
Through this, the current and voltage for all three phases 
shift in phase by 30° across the transformer. 



EXAMPLE 4.3 

The main substation at a mine contains a delta- 
delta connected transformer bank composed of three 
identical single-phase transformers. With rated volt- 
age applied, a 6,000-kW load at 0.8 lagging power 
factor is causing the transformer bank to be fully 
loaded. The rated primary and secondary voltages of 
each single-phase transformer are 36 kV and 7.2 kV, 
respectively. 

1. Find the capacity of each single-phase trans- 
former in the bank. 

2. What are the magnitudes of the primary and 
secondary currents in each single-phase transformer? 

3. What are the magnitudes of the primary and 
secondary line currents to and from the transformer 
bank? 

SOLUTION: The problem states that the trans- 
former bank is fully loaded by an average power, P T . 
Thus, the capacity or apparent power load, S T , of the 
bank is available from 

P T = St cos0, 



The capacity of each single-phase transformer, S P , 
one-third the total bank capacity, or 

fc>T = OOp 



IS 



and 



Sp = L 5 ™ = 2,500 kVA. 



If the transformer is assumed to be ideal, current 
and voltage in the primary or secondary are related 
to apparent power by 



Sp = 



Vplp. 



Hence, the primary and secondary currents in each 
single-phase transformer are 



V V pl 



2,500 
36 



= 69 A 



and 



±.± m m m ,; 



v. 



p2 



7.2 



It can be noted that these currents also correspond 
to the transformer turns ratio, a, which is 1/5 or 0.2. 
Because the transformer bank is delta-delta con- 
nected, the currents in each transformer are also the 
phase currents in the bank. Therefore, the line 
currents to and from the bank are, respectively, 



and 



I L = V3 I P 

I L1 = V3(69) = 120 A, 
I L2 = V3(347) = 601 A. 



BALANCED THREE-PHASE CIRCUIT ANALYSIS 

By definition, any element in one phase of a balanced 
(or symmetrical) system is duplicated in the other two 
phases. In other words, currents and voltages for the other 
phases are equal in magnitude but displaced symmetri- 
cally in phase position. Therefore, the analysis of voltage, 
current, impedance, and power in one phase can provide 
complete knowledge about the entire three-phase system. 
In addition, reactions between phases, such as phase 
curren ts, line-to-lin e voltage s, or line-to-line connected 
impedances, may be represented by an equivalent line or 
line-to-neutral value by using delta-wye transformations. 
The solution technique is called per-phase or single-phase 
analysis. The technique has wide application because 
almost all three-phase power systems that are operating 
normally are approximately balanced. 

As a simple demonstration of the concept, consider 
figure 4.13A, which illustrates a wye generator connected 
through line resistance to a wye-connected load. In figure 
4.13.B, one phase of this circuit is extracted, and here 

• V is one leg of the wye-connected source, 

• R is the line resistance per phase, 

• Z is a "single-leg" impedance of the wye-connected 
load, and 

• The unconnected points, n and n', are the neutrals 
of the source and load, respectively. 

For the balanced system, the vectorial sum of all three 
line currents is zero. Hence, the current between n and n' 
is zero, and the potential at n equals that at n'. Accord- 
ingly, the two neutrals can be joined as shown in figure 
4.13C This last diagram is the single-phase equivalent 



84 



circuit, single-phase diagram, or per-phase reduction of 
figure 4.13A. It should be noted that figure 4.13C is indeed 
a basic power circuit, similar to figure 4.1. 

Reduction of circuits containing delta-connected 
sources and loads is almost as easy, but one additional step 
is involved: the application of delta-wye transformation. 
Figure 4.14 demonstrates a simple example. Here, all 
sources and loads must be wye connected. The aim is to 
convert delta connections to wye using equation 4.9, and 
line-to-line voltages and phase currents to line-to-neutral 
and line, respectively. Thus, figure 4.14S is the per-phase 
equivalent of figure 4.14A. 

The simplified representation of the balanced three- 
phase circuit can now be analyzed, employing all the 
single-phase techniques previously discussed. When the 
solution is found, the three-phase parameters can be 
determined by reversing the reduction. This need only be 
performed when line-to-line or phase values are required; 
no changes are necessary with line-to-neutral and line 
parameters, as can be seen in figure 4.13. 



EXAMPLE 4.4 

A load has a balanced delta-connected imped- 
ance of 5 1 45° per leg. This load is connected through 
three balanced line impedances of 1 + jl Q to a 
three-phase source that has a line-to-line voltage of 
500 V. What is the magnitude of line current deliv- 
ered to the load? 

SOLUTION. This problem is basically the same as 
that for the circuit in figure 4.14, except here a line 
impedance exists between the source and the delta- 
connected load. As a per-phase solution is called for, 
the delta load must be transformed to an equivalent 
wye: 

Z A = 5|45^ n, 
Z Y = ^.5l^l.i.67|4610, 



or 



Z Y = 1.18 + jl.18 Q. 



The per-phase equivalent impedance as seen by the 
source is simply the sum of the line impedance and 
the equivalent wye impedance of the load, or 

Z eq = Z L + Z Y 

= 1 +jl + 1.18 + jl.18 

= 2.18 + J2.18 = 3.08 |45_1 Q. 

The magnitude of this impedance divided into the 
line-to-neutral voltage across any one phase yields 
the answer. 



Il " IZ 



Ln 



eql 

288 
3.08 



500 V3 
3.08 

= 94 A. 





|v an | = |v bn |=|v cn | = 

Rq = Rb = Re = R 

z a = z b = z c = z 




Figure 4.13.— Per-phase reduction of wye-to-wye system. 





Ra " Rb " Re " R 
Z a = Z b = Z c = Z 

Figure 4.14.— Per-phase reduction of delta-to-delta system. 



EXAMPLE 4.5 

A three-phase 200-hp induction motor has a 
full-load efficiency of 90%, power factor of 0.85 
lagging, and a rated terminal voltage of 950 V line- 
to-line. Find an equivalent delta-connected imped- 
ance for the motor when it is operating at full load 
under rated voltage. 

SOLUTION. Perhaps the best way to start this 
solution is to find the per-phase average power 
consumed by the motor under the stated conditions. 
The total average power input to the motor can be 
calculated from 



P T = 



(0.746) hp 



where hp is the motor horsepower and r\ is its 
efficiency. Thus, 



P T = 



(0.746X200) 
0.9 



= 165.8 kW. 



As single-phase analysis is desirable, the power 
consumed by each element of the equivalent wye- 
connected load is needed: 



165.8 



P -^ = 
^ p " 3 3 



= 55.3 kW 



85 



Equation 4.12 can now be used to find the line 
current to the motor, or 

Pp = V Ln I L cos0, 



II V t „ cos0 



55£60 
(950/VF) (0.85) 



= 118.5A. 



The line-to-neutral voltage divided by this line cur- 
rent is the magnitude of each leg of the equivalent 
wye-connected impedance for the motor, and the 
impedance angle is identical to the power-factor 
angle. Therefore, 

z Y = ^ \e 



950/V3 
118.5 



cos _1 0.85 



= 4.63131.8° fi; 



as the equivalent delta-connected impedance is re- 
quested, 

J Y 



Z A = 3Z V = 13.9|31.8 C 



or 



Z A = 11.8 + J7.3 a 



EXAMPLE 4.6 

A production shovel, operating at full load, uses 
1,200 kW at 0.9 lagging power factor with 3,750 V 
line-to-line at the machine. The shovel is supplied 
through a trailing cable that has an impedance of 0.04 
+ jO.03 fl per phase. If the voltage at the source side of 
the trailing cable is maintained constant, what is the 
voltage regulation at the machine? 

SOLUTION. The per-phase equivalent circuit for this 
problem is again similar to figure 4.145. The equivalent 
impedance of the shovel is not necessary, but the line 
currents and line-to-neutral voltage conditions for full 
load and no load are. Those for full load are 

p p = -2 = 400 kW, 

p 3 



Vfl = 



•M?T — 



' fe = 2,165 V line-to-neutral, 



400,000 
(2,165X0.9) 



= 205.3 A. 



The line-to-neutral voltage of the constant source, 
Vnl, can be found by computing the voltage drop 
across the trailing cable, V^, then adding it to the 
voltage at the machine. The full-load voltage at the 
machine can be assigned a zero phase angle. Thus, 
based on the given power factor, 

V FL = 2,165|0°_ V, 



I FL = 205.31 -25.84° A. 



The voltage drop across the trailing cable is 

V te = I FL Z te 

= (205.31 -25.84° ) (0.05| 36.9° ) 
= 10.271 11.06° V. 

The voltage at the constant source is 



= 10 + 



v FL 

j2 + 2,165 
= 2,175 + j2 = 2,175|0O^ V. 



The subscript for this voltage is used to signify that 
for these conditions it is the no-load voltage at the 
machine. In other words, under no load, line current 
through and the voltage drop across the trailing 
cable will be zero, and the voltage at the machine 
will be the same as at the source voltage. Conse- 
quently, the voltage regulation is 



VR = 



V FL 



2,175 - 2,165 



(100%) 



2,165 



(100%) = 0.5%. 



The solution may be difficult when there are delta- 
wye or wye-delta transformers in the system because of the 
30° phase shift of voltage and current between the wind- 
ings. In other words, a line-to-neutral transformer second- 
ary voltage transfers to a line-to-line primary value and 
vice versa. Obviously, when the three-phase system is not 
balanced, the per-phase reduction method cannot be ap- 
plied, and other more specialized techniques are required. 
These are discussed at the end of this chapter. 



One-Line and Three-Line Diagrams 

It is now apparent that practically all three-phase 
circuits consist of three conductors, three transformers, 
and so forth. When all these components are shown in a 
schematic, the drawing is called a three-phase diagram, as 
given in figure 4.15. Such diagrams can be extremely 
helpful, especially when the circuits are concentrated in a 
piece of machinery or power equipment, because they 
allow a complete view of component interconnections. 
They are imperative in manufacturing or troubleshooting. 

However, when the circuits are large, as in a complete 
mine power system, three-line diagrams are not only 



86 



cumbersome to draw but also exceedingly difficult to read 
and interpret. Furthermore, if the power system is nor- 
mally balanced, a three-line diagram is unnecessary since 
the system is always solved as a single-phase circuit 
composed of one line conductor and a neutral return. In 
these cases, the three-line diagram is replaced by a one- 
line diagram in which the interconnections or conductors 
between components are represented by single lines plus 
conventional symbols. This is a further simplification of 
the per-phase diagram because the completed circuit 
through the neutral is omitted. One-line diagrams are an 
invaluable tool in analysis, in designing new power sys- 
tems, or in modernizing existing ones. Furthermore, since 
circuit diagrams of coal mine power systems are required 
by Federal law (30 CFR 75, 77), a one-line diagram is the 
most practical way to comply. 

Figure 4.16 shows a one-line diagram designed to 
convey in concise form the significant information about 
the system shown in figure 4.15. In such diagrams it is 
usually implied that all information is per-phase, unless 
stated otherwise. Hence it is vital to remember that each 
device shown is actually installed in triplicate. The con- 
ventional notations are line or line-to-neutral impedance, 
line current, and line-to-neutral voltage. If line-to-line 
values are listed, they should be stated as such. Every line, 
symbol, figure, and letter has a definite meaning. Conse- 
quently, when a one-line diagram is constructed, specific 
conventions (1-2) 1 must be followed to ensure that the 
result will be complete, accurate, and correctly interpreted 
by anyone. The following guidelines are recommended. 

Relative geographic relationships for the power- 
system components should be maintained as far as prac- 
tical. The typical mine map provides an excellent layout 
guide for mining applications. Because of the shorthand 
form and definite meaning of every entry, duplication 
must be avoided. Standard numbers and symbols are 
mandatory, and those commonly used in mining are 
shown in figures 4.17 and 4.18 and tables 4.1 and 4.2. 
Many of the devices listed have not yet been covered but 
are included here for completeness. 

All known facts about the power system should be 
shown on the diagram, including 



on), 



• Apparatus ratings (volts, amperes, power, and so 

Ratios and taps of current and potential transformers, 

Power-transformer winding connections, 

Relay functions, and 

Size and type of conductors. 



The amount of information shown depends on the 
purpose of the diagram. For instance, if the diagram is to 
assist in studying the power flow to loads throughout the 
system (a load-flow study), the location of circuit breakers 
and relays is unimportant. However, for the solution of 
other power problems, complete knowledge of these de- 
vices can be mandatory. It is important that the one-line 
diagram contain only known facts; implications and 
guesses can lead to disastrous results. 

On many one-line diagrams, knowledge of future 
electrical plans is very helpful, and this information can 
be entered either diagrammatically or through explana- 
tory notes. Finally, the diagram should include correct 
title data so that the installation is clearly identified and 
cannot be confused with another or portion thereof. 




Z, =0.6 + j0.6 
Z 2 =0.07 + j 0.05 



Y-Y Transformer 
Figure 4.15.— Three-line diagram. 



O 



Line 
0.6 tj 0.6 



l:a 



Line 
0.07 + j 0.05 



1,385/277 

Z=0.1+j0.3 

referred to 

high side 




V=277V 



If machine, could 
be circle symbol 



1 Italicized numbers in parentheses refer to items in the list of references 
at the end of this chapter. 



Figure 4.16.— One-line diagram of circuit shown in figure 
4.15. 



87 



■«■ ■» 



Air circuit breaker, 
removable type DH 



Circuit breaker, 

nondrawout type 

(oil or vacuum) 



«-0^ 



Air circuit breaker, 
drawout type 



r\ 



Air circuit breaker, 
nondrawout type, 
series trip 



cHh 



«-o^ 



Magnetic starter 



Current-limiting 

breaker, 

drawout type 



-j&- 



Disconnecting fuse, 
nondrawout 



«d> 



Drawout fuse 



Disconnecting 
switch, nondrawout 



^^^ 



Disconnecting 
switch, drawout type 



-f^t 



Current 
transformer 



tt 



Potential 
transformer 



t> 



o-)|l 



Pothead 



Ground 



Surge arrestor 



Surge capacitor 



II 



> 



Battery 



Cable 
terminations 



Rectifier bank 



Reactor, 
nonmagnetic core 



Reactor, 
magnetic core 



Power 
transformer 



3-phase power 
transformer 
connected delta- wye 



Y 

Wye 

A 

Delta 



Figure 4.17.— Commonly used symbols for one-line electrical diagrams. 



88 



Overvoltage 



Undervoltage 



Overcurrent 



•'HW- 



Ground 
overcurrent 



> < 



Undercurrent 



« X ► 



Differential 
current 



■■i N X ► 

Differential 
ground current 



I A 

Balanced 
current 



Directional 
overcurrent 



■I 



-W— ► 



Directional 
ground overcurrent 



Directional 
power 



GP 



Gas pressure 



PW N 

Pilot wire 
(differential 

current) 



PW 

Pilot wire 
(directional 
comparison) 



4 Z ► 



Distance 



-i n z >: 

Ground distance 



^ 



1r+ 



Directional 
distance 



"HI Z ► 

Directional 
ground distance 



CC 

Carrier directional 

comparison 
(phase and control) 



J\ 



CC^ 



^Sh> 



Carrier phase 
comparison 



Synchro check 



4 S ► 

Auto 
synchronizing 



d fc 

< A > 



Phase balance 



-0— ► 



Phase rotation 



4 r ^u ► 



Overfrequency 



« T ► 



Overtemperature 



Note: 

For directional relays, arrow 
points toward fault that will 
cause tripping 



Figure 4.17.— Commonly used symbols for one-line electrical diagrams— Con. 



89 



4 



1 



Wye, neutral ground 



<^ 



Zig-zag ground 



j/ 



CT 




' Meter 



Current transformer 
with ammeter; 
letter indicates 
instrument type 



Hr^ 

■&S-45 



-0 

Relay 



Relays connected 

to CT's and PT's. 

Number indicates 

relay type function 



1 



Ground 



RELAY FUNCTIONS 



± 



Overcurrent 



Differential 



O 



Induction motor 
or general source 



O 



Synchronous 
motor 



Instrument transfer 
switch. Letter 
indicates type 



< > 



Dummy circuit 

breaker. 
Removable type 



-< >- 



Future breaker 

position. 
Removable type 



-WW 



Resistor 



Figure 4.18.— Symbols for relay functions. 



90 



Table 4.1. —IEEE device numbers and functions (1) 



Device 



Function 



1 Master element. 

2 Time-delay starting or closing relay. 

3 Checking or interlocking relay. 

4 Master contactor. 

5 Stopping device. 

6 Starting circuit breaker. 

7 Anode circuit breaker. 

8 Control power-disconnecting device. 

9 Reversing device. 

10 Unit sequence switch. 

1 1 Reserved for future application. 

12 Overspeed device. 

13 Synchronous-speed device. 

14 Underspeed device. 

15 Speed, or frequency matching device. 

16 Reserved for future application. 

17 Shunting or discharge switch. 

18 Accelerating or. 

19 Starting to running transition contactor. 

20 Electrically operated valve. 

21 Distance relay. 

22 Equalizer circuit breaker. 

23 Temperature control device. 

24 Reserved for future application. 

25 Synchronizing or synchronism check. 

26 Apparatus thermal device. 

27 Undervoltage relay. 

28 Reserved for future application. 

29 Isolating contactor. 

30 Annunciator relay. 

31 Separate excitation device. 

32 Directional power relay. 

33 Position switch. 

34 Motor-operated sequence switch. 

35 Brush-operating or slip-ring short-circuiting device. 

36 Polarity device. 

37 Undercurrent or underpower relay. 

38 Bearing protective device. 

39 Reserved for future application. 

40 Field relay. 

41 Field circuit breaker. 

42 Running circuit breaker. 

43 Manual transfer or selector device. 

44 Unit sequence starting relay. 

45 Reserved for future application. 

46 Reverse-phase-balance current relay. 

47 Phase-sequence voltage relay. 

48 Incomplete sequence relay. 

49 Machine or transformer thermal relay. 

50 Instantaneous overcurrent or rate-of-rise relay. 



Device 



Function 



51 Time overcurrent relay — ac. 

52 Circuit breaker— ac. 

53 Exciter or dc generator relay. 

54 High-speed dc circuit breaker. 

55 Power-factor relay. 

56 Field application relay. 

57.. Short-circuiting or grounding device. 

58 Power rectifier misfire relay. 

59 Overvoltage relay. 

60 Voltage balance relay. 

61 Current balance relay. 

62 Time delay stopping or opening relay. 

63 Liquid or gas pressure level or flow relay. 

64 Ground protective relay. 

65 Governor. 

66 Notching or jogging device. 

67 ac directional overcurrent relay. 

68 Blocking relay. 

69 Permissive control device. 

70 Electrically operated rheostat. 

71 Reserved for future application. 

72 Circuit breaker — dc. 

73 Load resistor contactor. 

74 Alarm relay. 

75 Position-changing mechanism. 

76 Overcurrent relay — dc. 

77 Pulse transmitter. 

78 Phase-angle measuring or out-of-step protective relay. 

79 Reclosing relay — ac. 

80 Reserved for future application. 

81 Frequency relay. 

82 Reclosing relay— dc. 

83 Automatic selective control or transfer relay. 

84 Operating mechanism. 

85 Carrier or pilot-wire receiver relay. 

86 Locking-out relay. 

87 Differential protective relay. 

88 Auxiliary motor or motor generator. 

89 Line switch. 

90 Regulating device. 

91 Voltage directional relay. 

92 Voltage and power directional relay. 

93 Field-changing contactor. 

94 Tripping or trip-free relay. 

95 Reserved for future application. 

96 Do. 

97 Do. 

98 Do. 

99 Do. 



Table 4.2.— Device numbers and letters common to mining (2) 



Device 



Function 



1 Control switch. 

3 Plus interlock relay. 

37 Ground-continuity check undercurrent relay. 

49 D Diode thermal relay. 

49 GR Grounding resistor thermal relay. 

49 T Transformer thermal relay. 

50 Instantaneous overcurrent relay— ac. 

50 G Instantaneous overcurrent relay — dc — connected in 

ground wire. 

50 N Instantaneous overcurrent relay— ac — connected to 

neutral. 

50 Z Instantaneous current-balance relay — ac— zero sequence. 

51 Time delay overcurrent relay — ac. 

51 N Time-delay overcurrent relay— ac — connected to neutral. 

51 Z Time-delay current-balance relay— ac— zero sequence. 

52 Circuit breaker — ac. 

59 GR Ground protective relay — dc — unbalance relay. 

72 Circuit breaker — dc. 

76 Overcurrent relay — dc. 

A Ammeter. 

D Demand meter. 

GD Ground detector. 

PF Power factor. 

V Voltmeter. 

VA Volt-ammeter. 

VAR Varmeter. 

W Wattmeter. 

WH Watthour meter. 



The remainder of the text will employ one-line and 
per-phase diagrams almost exclusively. The main thing to 
remember is that practically every item represents three 
identical or corresponding items in the actual system. 
Even when the normally balanced system becomes unbal- 
anced through component failures, the same diagram is 
used, the only change being the notation for the specific 
failure. 



Circuits Containing Transformers 

As previously stated, solving a balanced three-phase 
system problem by per-phase analysis is as simple as the 
single-phase techniques covered in chapters 2 and 3. 
However, the solution is not so clear when delta-wye or 
wye-delta transformers are involved. Perhaps the simplest 
way to demonstrate the approach is first to illustrate a 
problem solution where the per-phase reduction is of a 
straightforward wye-wye transformer, then change to a 
delta-delta or delta-wye transformer and show the compli- 
cations that might arise. 



91 



EXAMPLE 4.7 

Consider figure 4.19, which shows a one-line 
diagram of a substation supplying power through 
about 1 mile of overhead line (power conductors on 
poles) to a three-phase wye-wye transformer bank, 
then through a trailing cable to a three-phase induc- 
tion motor. The motor consumes an average three- 
phase power of 150 kW, operating at 0.8 lagging 
power factor. The problem is to find the rms voltage 
needed at the substation output to provide the rated 
motor terminal voltage of 480 V line-to-line. 

A three-phase diagram of the circuit is shown in 
figure 4.20 for reference. The first step in the solu- 
tion is usually to develop a per-phase circuit as 
shown in figure 4.21. Although the solution could be 



Substation 



Load 
center 



Utility 






Overhead line 
Z = 0.6+jO6 



Trailing cable 
' Z=0.07 + j0.5 



c (t) 



-ir 



o 



-»- 



^a 



1,385 /277 V 

Z = 0.1+jO3 

referred to high side 



50 kW at 

0.8 lagging pf 

277 V line-to-neutral 



Figure 4.19.— One-line diagram for example 4.7. 




= Z 2 = 0.07 +j 0.05 



Transformer 



Figure 4.20.— Three-phase diagram of figure 4.19. 



5M 

(r) Zol=0.6+j0.6 Z X p=0.1+j0.3 Z tc = 007 +j 0.05 

-VWV V-r VWV 1 r—n WW- 



Vt 



V, 



V L =277V 



Figure 4.21.— Per-phase diagram of figure 4.19. 



performed directly from the one-line diagram, the 
per-phase diagram allows direct application of 
single-phase techniques. The following should be 
noted in figure 4.21. 

• The line and trailing-cable conductor imped- 
ances are now illustrated as circuit elements. 

• Only one phase of the transformer bank is 
shown, represented as an impedance in series with 
the primary of an ideal transformer. The trans- 
former turns ratio is computed from 



a = 



Vg _ 277 1 
Vj " 1,385 " 5 ' 



where these rated voltages are line-to-neutral rms. 
• The induction motor is represented by its 
single-phase equivalent, P , where 



150 



Fp " 3 3 



- 50 kW. 



The solution can now follow a stepwise process. 

1. Assume the motor terminal voltage, V L , is the 
rated 277-V rms line-to-neutral (480/V3). 

2. Compute motor line-current magnitude, I L , 
using 



or 



P P = V L I L (load pf) 
P P 50,000 



II V L (load pf) (277X0.8) 



= 225.5 A. 



If the motor terminal voltage is taken as the refer- 
ence phasor, this current has a phase angle deter- 
mined by the load power factor. Therefore the motor 
current phasor is 

I L = It.I -cos-^O.S) = 2261 -36.9° A. 

3. I L is then employed to find the voltage drop 
across the per-phase equivalent of the trailing cable, 
Z^. When this is added to the motor terminal 
voltage, the voltage at the ideal transformer second- 
ary, V 8 , is 

V 8 = I L Z te + V L 

or V 8 = (226| -36.9° X0.09| 35.5° ) + 277|0°_ 
= 20.301 -1.4° + 277|0^ 
= 20.29 -J0.50+ 277 
= 297.29 -j 0.50 + 297| -0.1° V. 

4. For the ideal transformer with a turns ratio, a, 
of 1/5, the voltage across the primary is 



v -i 

p a 
or V P = 5(2971 -0.1° ) = I486] -0.1° V, 



92 



with the primary current being 

I P = ai L 
or I P = (|X226| -36.9 ) = 45| -36.9° A. 

Notice that there is no change in current or voltage 
phase angle across the transformer. 

5. This primary current can now be used to find 
the voltage drop resulting from transformer and 
overhead-line impedances. Summing this voltage 
drop with the transformer primary voltage gives the 
desired answer to the problem, the substation out- 
put voltage, V t : 

v t = I P (Z OL + Z xp ) + Vp, 

where Z OL , Z xp = overhead-line and transformer 



Then, 



impedances, respectively. 



V t = 451 - 36.9° (0.6 +J0.6 + 0.1 +.J0.03) 
+ 1,486|-0.1° 



or V t = 45| -36.9° (0.94 | 41.99° ) + 1,486| -0.1° 
= 42.39|5J^ + 1,486 | -0.1° 
= l,528| +0.04° V. 

Because the analysis is per phase, the result is 
obviously line-to-neutral voltage. If line-to-line val- 
ues are required, the above answer need only be 
multipled by V3. It is interesting that in this exam- 
ple the phase angle of the substation voltage is 
practically the same as that at the motor. This may 
not be the case in actual mine power systems. 



When the transformer is delta-delta connected, problem 
solutions are practically the same as in example 4.7. The 
one-line diagram of figure 4.22 provides an illustration. 
When the system is represented per phase (fig. 4.23), the only 
additional concern is delta-wye transformation of the trans- 
former impedance; then the solution proceeds as before. 
However, the process may not be as simple when delta-wye or 
wye-delta transformer connections are involved. 

Consider the one-line diagram in figure 4.24 that 
shows a delta-wye transformer supplying the same motor 
as that shown in figure 4.19. Figure 4.25 illustrates one 
leg of the three-phase transformer. From this, it can be 
seen that secondary line currents appear as phase cur- 
rents on the primary, and line-to-neutral secondary volt- 
ages become line-to-line primary voltages. In other words 
(fig. 4.25), for primary voltage in terms of the secondary 
(the ideal transformer with turns ratio, a, only): 



V AR = 



_an | 30 c 

a — 



(4.18) 



where V_ab = line-to-line primary voltage, V, 
and V an = line-to-neutral secondary voltage, V, 

and for primary and secondary current, 

I AB = a IJ301, 

where I^p = primary phase current, A, 
and I a = secondary line current, A. 



(4.19) 



Substation 



Z = 0.6+j0.6 ] £ Z=0.07 + j0.05 

*-; " r» — < s — • — —- 



:t) 



A J ""A 

Bank of 3-10 XFMRS, 
each 2.400/480V 

Z=0.3+j0.9 
referred to high side 



o 



150 kW total at 

0.8 lagging pf 

480 V line-to-line 



Figure 4.22.— One-line diagram with delta-delta transformer. 



^-xp ^-xp'^ •-''' 
(t) Z 0L =0.6+j0.6 =Ql+j0.3 Z tc = 0.07 +j 0.05 
-VWV r*7 WvV 1 i — r*7 VWv- 




V L =277V 



Figure 4.23.— Per-phase diagram of figure 4.22. 



Load 
Substation center 

Utility -\ r Overhead line > r 



Z = 0.6 + j0.6 



Trailing cable 
Z =0.07 tj 0.5 



(t) 



-O 



"la 



1,385 /277 V 
Z=0.1+j0.3 
referred to high side 



50 kW at 

0.8 lagging pf 

277 V line-to-neutral 



Figure 4.24.— One-line diagram with delta-wye transformer. 




Primary Secondary 

Figure 4.25.— One leg of three-phase transformer from 
figure 4.24. 



However, when performing balanced three-phase 
analysis, the parameters of interest are line-to-neutral 
voltages and line currents. Thus, to continue the analysis 
in a fashion similar to that used in example 4.7 (the 
wye-wye transformer), V^ and Ij^ must be converted to 
their respective per-phase equivalents. Recalling that 






V ab = V3 V„J+30°, 



(4.20) 



and applying this concept to equation 4.18, the primary 
line-to-neutral voltage, V An , is 



_ V 

V - *£. 
VAn_ aV3- 



(4.21) 



93 



Employing equation 4.76 to convert equation 4.19, pri- 
mary line current, I A , is 



voltamperes. The mathematical interrelations of the bases 
are as follows: 



I A = a L 



(4.22) 



Equations 4.21 and 4.22 simply state that the phase shifts 
that occur across delta-wye or wye-delta-connected trans- 
formers do not interfere with the analysis when this is 
performed per phase. Analysis can be enhanced by chang- 
ing the delta primary or secondary to an equivalent wye 
connection, thus enabling the construction of a per-phase 
diagram for the entire system. 

Concerning the actual per-phase analysis, it has been 
shown that the three-phase circuit is reduced by a process 
no more difficult than the single-phase work covered in 
chapters 2 and 3. The next section will present a technique 
that further simplifies power-system analysis. 



PER-UNIT SYSTEM 

Problems related to electrical circuits should be solved 
in terms of volts, amperes, voltamperes and ohms. The 
answers to mine power problems, and indeed any electrical 
problem, almost always require these terms, but in the 
process of computations it is often more convenient to 
express these quantities in percent or per-unit (pu), re- 
ferred to some arbitrarily chosen base. For example, if a 
base voltage of 100 kV is selected, voltages of 90, 120, and 
125 kV have percent representations of 90%, 120%, and 
125%, or per-unit values of 0.9, 1.2, and 1.25 pu, respec- 
tively. Both percent and per-unit values express a ratio of 
a specific quantity to the base quantity. Per-unit is given 
as a decimal, whereas the ratio in percent is 100 times the 
per-unit value. These expressions, especially per-unit, are 
becoming standard for equipment specifications. 

There is a definite advantage in using per-unit nota- 
tion over percent. Per-unit multiplication or division 
yields a result in per-unit. However, the product of two 
percent quantities must be divided by 100 to obtain a 
percent answer. For example, if two quantities are both 0.9 
pu or 90%, then 



and 
but 



(0.9 puX0.9 pu) = 0.81 pu 

(90%) (90%) * 8,100% 
8,100 „ 



(90%X90%) = 



100 



Consequently, per-unit notation will be employed almost 
exclusively, the only exception being where conventions 
dictate otherwise. 

Voltage, current, voltamperes, and impedance are 
obviously interrelated for any specific circuit or system. As 
a result, the selection of any two determines the base 
values for the remaining two. For example, if the current 
and voltage bases are specific, the base impedance and 
base voltamperes can be found. Since three-phase circuits 
are usually solved as a single line with a neutral, base 
quantities in the per-unit system are line current, line- 
to-neutral voltage, per-phase impedance, and per-phase 



v b = i b z b , 



KVA b = kV b I b , 



T kVA b 
b " kV„ 



V b (kV^l.OOO 

by. — - — 



Iv 



kVA, 



(kVb) 2 
MVA h 



(4.23) 



(4.24) 



(4.25) 



(4.26) 



where V b = base line-to-neutral voltage, V, 
kV b = base line-to-neutral voltage, kV, 
kVA b = base per-phase voltamperes, kVA, 
MVA b = base per-phase voltamperes, MVA, 
I b = base line current, A, 
and Z b = base per-phase impedance, fi. 

All these formulas are adaptations of the fundamental 
Ohm's law and power material; the last three are ex- 
pressed in kilovolts and kilovoltamperes because of the 
levels normally found in power systems. It should be 
remembered that line-to-line voltages and total power 
(kilovoltamperes or megavoltamperes) are customarily 
specified; these must be changed to line-to-neutral volt- 
ages (dividing by V3) and per-phase power (dividing by 3) 
before equations 4.23 through 4.26 can be applied. 

To apply the per-unit system to power problems, base 
values for kilovoltamperes and kilovolts are normally 
selected first in order to minimize calculations as much as 
possible. Base values for impedance and current are then 
found. Next, all the actual voltages, currents, impedances, 
and powers in the power system or system segment are 
expressed as a ratio to the base quantities; these are the 
per-unit quantities. Problems are then solved in per-unit, 
with the answers converted back to actual parameters. 
The actual values and per-unit quantities are related by 



Za - Z pu Z b , 


(4.27) 


Ia = Ipu lb, 


(4.28) 


v A = v pu v b , 


(4.29) 


VA A = VAp U VA b) 


(4.30) 



where Z A , I A , V A , VA A = actual values of impedance, 

current, voltage, and power, re- 
spectively, fi, A, V, VA, 

and Z pu , Ip U , V pu , VAp U = per-unit values of impedance, 

current, voltage, and power, re- 
spectively, pufi, puA, puV, 
puVA. 

It is important to note that all impedances in a problem 
are referenced to the same base impedance, whether they 
are pure resistance or pure reactance. The same holds for 
all average, reactive, or apparent powers, which are refer- 
enced to the base kilovoltamperes. 



94 



Often, per-unit impedances or percent impedances of a 
system component have already been assigned to a base 
referenced to the component or power-system segment in 
which that component is located. These impedances can be 
changed to another base impedance by 



= Z Due kVA b /kV^ 2 
pu kVA„ IkV, 



(4.31) 



where Z 



ie = per-unit impedance of specified component, 

puft, 
kV e , kVA e = base kV and kVA used to reference Z pue , V, 

VA, 
kV b , KVA b = base kV and kVA to which new per-unit 

impedance is to be referenced, V, VA, 
and Z pu = new per-unit impedance referenced to kV b 

and kVA b , pu ft. 



Transformer Impedance 

Transformers are the most common devices in power 
systems where the component impedance is referenced to 
the rated power and voltage of the component. Convention- 
ally, percent impedance is specified, but this can be 
converted to per-unit simply by dividing by 100. 

A major advantage of using per-unit computations is 
seen when circuits are connected through transformers. 
Through the proper selection of voltage bases, the per-unit 
impedance of the transformer is the same no matter which 
winding is used. Consequently, if exciting and magnetiz- 
ing currents are ignored, as they often can be in power 
systems, the transformers become a simple series imped- 
ance in per-unit calculations. In other words, the ideal 
transformer is not needed in the equivalent circuit. Exam- 
ple 4.8 explores this advantage. 



EXAMPLE 4.8 

Consider a 750-kVA power-center transformer, 
the approximate per-phase equivalent circuit for 
which is shown in figure 4.26. The per-phase ratings 
are 250 kVA, 5,000/1,000 V, and 5-ft reactance re- 
ferred to the high side. Following convention, the 
base power kVA bl is 250 kVA, and the base voltage 
for the high side kV bl is 5 kV. One kilovolt is 
selected as the low-side base voltage, kV b2 . With 
these, the high-side base quantities can be calcu- 
lated using equations 4.23 through 4.26: 

kVA bl = 250 kVA, 



kV b 

Ik, 



and 



= 5kV, 
= 50 A, 



Zb! = 100 ft. 
The per-unit impedance of the transformer is thus 



rj Z^l 



100 



= 0.05 pu. 



Now consider the actual transformer impedance as 
it would appear referred to the low side, as in figure 
4.27. 



/N 



V, 



or 



Z A2 - Z A1 \^J - Z A1 ( y * 



The base quantities on the low side are 

kVA bl = 250 kVA, 
kV b2 = 1 kV, 



I, 



and 



= 250 A, 
= 4 ft. 



Notice that the base power does not change and the 
low-side base voltage defines base current and im- 
pedance. The per-unit transformer impedance as 
seen from the low side is 



-7 ^A2 

V = -7 

^b2 



0.2 



= 0.05 pu. 



Therefore, the per-unit impedance of the trans- 
former is the same, regardless of the side it is viewed 
from, and the per-unit equivalent circuit is simply 
the series circuit shown in figure 4.28. Here, the 
input and output voltages are now expressed in 
per-unit since the transformer is operating at rated 
voltage. 



Z A1 =j5n 



V,= 5,OOOV 




V 2 = 1,000 V 



Primary Secondary 



Figure 4.26.— Approximate per-phase equivalent circuit for 
750-kVA load-center transformer; impedance referred to high 
side. 



Z A2 =j0.2il 



V,= 5,000 V 




V 2 = 1,000 V 



Figure 4.27.— Transformer of figure 4.26 with impedance 
referred to low side. 



Z=j0.05pu 



V, = 1.0pu 



V 2 = 1.0pu 



Figure 4.28.— Simplified equivalent circuit of transformer 
expressed in per-unit. 



95 



Three-Winding Transformers 

In chapter 3 and to this point in chapter 4, equivalent 
circuits have been shown only for two-winding transform- 
ers, those having only one primary and one secondary 
winding. However, many transformers in mine power 
systems have three windings, with the third winding 
termed the tertiary or second secondary. These include 
power-center transformers supplying two different utiliza- 
tion voltages, such as 950 and 550 Vac to face equipment 
or 550 and 250 Vdc to machinery. The latter case not only 
uses a three-winding transformer but also three-phase 
rectification, which will be described in chapter 5. 

Both the primary and secondary windings of the 
two-winding transformer have the same kilovoltampere 
capacity or rating, but the three windings of a three- 
winding transformer may have different kilovoltampere 
ratings. The impedance of each winding may be given in 
percent or per-unit, based on each winding rating. The 
three impedances can also be measured by the following 
short-circuit test, where rated voltage is applied to the 
primary for Z ps and Z pt and to the secondary for Z 8t : 

Z ps = leakage impedance measured in primary (or 
first winding), with secondary (or second wind- 
ing) short-circuited and tertiary (or third wind- 
ing) open, Q, 

Z pt = leakage impedance measured in primary with 
tertiary short-circuited and secondary open, Q, 

Z st = leakage impedance measured in secondary 
with tertiary short-circuited and primary open, 



The impedances of the primary, secondary, and tertiary 
windings are found from 



or 



where Z p , Z s , Z t = impedances of primary, secondary, and 
tertiary, respectively, Q. 



they must have the same kilovoltampere base. Further, 
voltage bases for the circuits connected through the trans- 
former must have the same ratios as the turns ratio of the 
transformer windings; that is, primary to secondary, pri- 
mary to tertiary, which are actually the same as the ratios 
of the related winding voltages. 

Per-Unit Method in System Analysis 

As mentioned earlier, the use of per-unit equivalents 
in the analysis of power-system problems can greatly 
simplify the work involved, especially when the system 
contains transformers and different voltage levels. How- 
ever, as per-unit calculations require the change of famil- 
iar parameters (ohms, volts, amperes, and so on) into 
values representing a ratio, this advantage is often diffi- 
cult to comprehend. Example 4.9 will illustrate the per- 
unit method of analysis using the one-line diagram pro- 
vided in figure 4.30, which could represent a mine power 
system in the early stages of development. All power levels 
listed are given per-phase; those shown for the mining 
equipment represent consumption. The voltages are all 
line-to-neutral. 



EXAMPLE 4.9 

The information required could be the voltage or 
current level at any point. Regardless, solution by 
the per-unit method first requires translation of the 
impedance of all components to the same base. The 
base selection is arbitrary, but for convenience, the 
largest kilovoltampere capacity of a system compo- 
nent is usually taken. In this case a good base would 



Z p8 = Z p + Z s , 




(4.32a) 




Z pt = Z p + z t) 




(4.326) 




Z 8t = z 8 + z t , 




(4.32c) 


Po 3 C C OS 


Zp = 2 (Z ps + Z pt - 


Z 8t ), 


(4.33a) 


*5t C-<- 


Z 8 = 2 (Z ps + Z 8t - 


Zpt), 


(4.336) 


6 t 


Z t = 2 ( Z pt + Z at - 


Zps), 


(4.33c) 


Hine symbol 



pa- 



-A/WV- 



^vwv- 



-oS 



z» 

-AAMr 



-ot 



Neutral 



Equivalent circuit 



^n 



Figure 4.29.— Approximate equivalent circuit of three- 
winding transformer expressed in per-unit. 



In equations 4.32 and 4.33, all impedances (Z p8 , Z pt , Z st ) 
must be referred to the primary winding voltage. If Z 8t is 
obtained from the described measurements, the imped- 
ance is referred to the secondary- winding voltage, hence it 
must be transferred. 

The approximate per-phase equivalent circuit for a 
three-winding transformer with the winding impedances 
of Z p , Z 8 , and Z t is provided in figure 4.29. Magnetizing and 
exciting currents are ignored. The terminals p, s, and t are 
the primary, secondary, and tertiary connections; the 
common point is unrelated to the system neutral. The 
three impedances must be in the per-unit system, as was 
the case for the equivalent circuit in figure 4.28. Hence 



Substation center 

Feeder cable 



^ y H Z=0.13 + j0.06 H 

^^il — — it 



Trailing cable Miner 

Load p O.Q3 + j 0.01 p = 57 k W 

~^WQ=47kvar 
V=320V 



1,000 kVA 

40/7.2 kV 

Z = 7% 



150 kVA 

7.2/350kV 

Z = 4.5% 



^ 



Trailing cable 

Z = 0.1+ jO.02 _ Shuttle 

• Q 



Bus 



P=4kW 
Q=5kvar 



Figure 4.30.— One-line diagram of small mine power system. 



96 



be 1,000 kVA, corresponding to the per-phase capac- 
ity of the substation. But two base parameters are 
needed in order to define the four base quantities; as 
the nominal voltage for each voltage level can be or 
can approach a constant, the system voltages are an 
excellent choice for the second base parameter. For 
figure 4.30, these would be the line-to-neutral volt- 
ages of 40 kV at the utility, 7.2 kV at mine distri- 
bution, and 350 V at mine utilization. Note that the 
system voltages are usually given as line-to-line in 
one-line diagrams, so they must be changed to 
line-to-neutral values to employ the formulas given 
here. In any event, base voltages must correspond 
with the turns ratio of any interconnecting trans- 
former. The ones selected do. 

With base kilovoltamperes and base kilovolts 
specified, the base quantities can be calculated 
using equations 4.23 through 4.26. 



1. For the utility: 

kVA b = 1,000 kVA, 

kV bl = 40 kV, 
_kV^_ i^oo 

J-bl - UT7 - An - S.O J\, 



kV bl 



40 



Zbi J^l^OO m (40^,000 m 1600 



kVA, 



1,000 



2. For mine distribution: 

kVA b = 1,000 kVA, 
kV b2 = 7.2 kV, 



Zb2 ~ 



2 ~ 7.2 
(7.2)2 1,000 



1,000 



= 52 fl. 



3. For mine utilization: 

kVA b = 1,000 kVA, 

kV b3 = 0.35 kV, 

1,000 „ „_, . 
Ik, = -:h^ = 2,857 A, 



Zb3 - 



0.35 

(0.35) 2 1,000 
1,000 



= 0.12 Q. 



The per-unit representations for all components 
of the mine system can now be found, and the 
needed formulas are equations 4.27 through 4.31. 



1. For the substation: percent reactance is 7% or 
0.07 pu based on the transformer rated kilovoltam- 
peres, referred to the high side, 40 kV. 



z = Z DUe kVA b £V^ 2 



kVA e \kV b 

where Z pue = J0.07 pu, 

kVA b = kVA e = 1,000 kVA, 
kV e = kV b = 40 kV; 
thus, Z pue = J0.07 pu. 

2. For the feeder cable: actual impedance is 
given, 

Z A = Ra + JX A = 0.13 + J0.06 Q, 
Za _ 0.13 + J0.06 

'b2 



and 



Zp U - 7 



52 



= (0.0025 + jO.0012) pu. 



3. For the load center: percent reactance is 4.5% 
or 0.045 pu based on the transformer rated kilovolt- 
amperes, referred to the high side, 7.20 kV. 



_ Z Due kVA b / kV e , 2 
pu kVA„ IkVu 



where 



and 



Zpu = 



Z pue = J0.045 pu, 

kVA b = 1,000 kVA, 

kVA e = 150 kVA, 

kV b2 = kV e = 7.2 kV, 
(j0.045Xl,000) 



150 



= jO.3 pu. 



4. For the trailing cables: actual impedances are 
again given. 

Miner. Z A = 0.03 + J0.01 fi, 



Za _ 0.03 + jO.Ol 
pu " Z b3 " 0.12 



= (0.25 + jO.083) pu. 



Shuttle car. Z A = 0.1 + J0.02 12, 



7 - --^- - 

pu ~ Z ~ 

^b3 



0.1 + jO.02 
0.12 



= (0.833 + jO.167) pu. 



97 



5. For the machines: consumption is given in 
terms of average and reactive power. 

Miner. P = 57 kW, Q = 45 kvar 

or kVA A = (57 + j45) kVA; 

thus, kV V - kYA - - (5 ^^ ) 



kVA, 



1,000 



= (0.057 + jO.045) pu. 



Shuttle car. kVA A = (4 + j5) kVA, 

hence, kVA„ u = (0.004 + jO.005) pu. 

At this point, the entire system of figure 4.30 may be 
redrawn into the impedance diagram in figure 4.31. 

Figure 4.31, when compared with a per-phase 
diagram in the impedance domain such as figure 
4.21, illustrates the advantage that simplified per- 
unit computations lend to power-system analysis. 
The circuit shown is merely a series-parallel ar- 
rangement of basic electrical elements, and obvi- 
ously it may be further simplified if desired, say into 
an equivalent per-unit impedance. This is only one 
example; an actual appreciation of power-system 
analysis by per-unit techniques can come only 
through experience. 

The impedance diagram can be used for the 
solution of most power problems. Suppose currents 
under normal operation are desired. One method is 
to apply known voltages at system points and calcu- 
late the resulting currents and voltages. For in- 
stance, suppose the line-to-neutral at the miner is 
320 V (about 555 V line-to-line). The per-unit equiv- 
alent of this is 



tv pu . ^ 



kV 



b3 



0.32 
0.35 



= 0.91 pu. 



The line current through the miner trailing cable is 
then 



or 



_ /kVA^ 
^ u \ kV. 



(0.057 - jO.045) nntin . nnAn 
^ = 091 = 0063 " j0049 pu - 



The conjugate of power is employed because voltage 
is implied as the reference phasor following the 
conventions stated earlier. The process is then con- 
tinued through the entire system. When the desired 
per-unit values are obtained, they are simply con- 
verted to actual values. Considering the current in 
the miner trailing cable, the actual line current is 

Ia = Ipu Ib3 

or I A = (0.063 - J0.049X2857) = 180 - J141.3A 
or I A = 229 1 -38.1° A. 



(0.25 + j 0.083) pu 



Utility 



j0.07pu j0.3pu 



(0.0025 +j 0.0012) 



Miner load kVA 
"> (0.057 + jO045)pu 



(0.833 + j 0. 167 )p U 



Shuttle car kVA 
' (0.004 + j 0.005) pu 



Figure 4.31.— Impedance diagram of system in figure 4.30, 
expressed in per-unit on a 1,000-kVA base. 



UNBALANCED THREE-PHASE CIRCUITS 

The solution of balanced three-phase circuit problems 
is usually accomplished by converting the constants, cur- 
rents, and voltages to per-phase values. Because symme- 
try determines the magnitude and phase position of all 
currents and voltages, actions occurring between phases 
can be represented by equivalent impedances. Further- 
more, currents and voltages for the other phases are equal 
in magnitude to those in the per-phase solution but are 
simply displaced symmetrically in phase position. This is 
extremely important because normally operating three- 
phase power systems can usually be approximated as 
balanced. However, the solution of unbalanced three-phase 
circuits or balanced circuits with unbalanced loads does 
not permit the same simplification. 

Mine power systems are designed to have a high degree 
of reliability and therefore to operate in a balanced mode. 
But at times, equipment failures and unintentional or inten- 
tional disturbances from outside sources can result in an 
unbalanced operation. The most common sites for mine 
power-system unbalance are equipment trailing cables. The 
consequence of unbalance is abnormal currents and volt- 
ages, and if safeguards are not designed into the system to 
protect against these anomalies, the safety of personnel as 
well as equipment can be compromised. The protective 
circuitry within the mine power system serves as the safety 
valve for such hazardous m alfunctions. 

Power-system unbalance can occur either from open 
circuits or from faults. A fault occurs whenever electricity 
strays from its proper path. Faults can be visualized as the 
connecting together of two or more conductors that nor- 
mally operate with a voltage between them. The connec- 
tion that creates a fault can be from physical contact or an 
arc caused by current flow through a gaseous medium. A 
short circuit is one type of fault. Currents in the power 
system resulting from an open circuit or a fault can be 
exceedingly large. 

An overload is not a fault. The term overload merely 
implies that currents exceed those for which the power 
system was designed. Such currents are usually much 
smaller than fault currents. Nevertheless, overloads can 
create equipment failures by exceeding the thermal design 
limits of the system. If not corrected, the overload can 
result in a hazard to personnel. However, such problems 
occur only with unattended overload operation for an 
extended time period, whereas faults can create an unsafe 
condition almost instantly. 

Both circuit breakers and fuses are used to protect 
circuits from excessive current flow, be it a result of 
faulting or overloading. The circuit-interrupting operation 
consists of parting a pair of contacts, and since an arc is 



98 



drawn between the contacts, the process must also extin- 
guish the arc. The interruption is handled mechanically in 
the circuit breaker, and the excessive current is monitored 
electrically or thermally. Fuse operation is based on sim- 
ple thermal operation. The fusible element is responsive to 
the heat of an overload or fault current and melts open. 
The fuse jacket is employed to extinguish the subsequent 
arc. A complete discussion of interrupting devices and the 
associated protective circuitry is presented in chapter 9. 

Fault Types 

The fault type often seen in literature is called a 
bolted fault, which can be described as a zero-resistance 
short circuit between two or more conductors. In reality 
most faults are not dead shorts but have some finite value 
of resistance. 

Faults may be classed as permanent or temporary. A 
permanent fault is one where equipment operation is 
impossible and repairs are mandatory. A temporary fault 
is intermittent in nature. For instance, two closely spaced 
overhead conductors may cause trouble only on windy 
days, when they can be forced into contact or close prox- 
imity by the wind. 

A very sinister fault type is the arcing fault. This 
condition is now believed to be the most common fault. 
When two conductors of different potential are in very 
close proximity, the intervening space between them can 
be considered as a spark gap. If the two conductors are part 
of an ac power system, the insulating material between 
the conductors may break down when the sinusoidal 
waveform reaches a certain value. Fault current will then 
flow. The potential drop across the conducting gap, which 
is actually an impedance, remains at a nearly constant 
level. It is this energy source, releasing terrific quantities 
of heat, that causes the devastation that is typical of an 
arcing fault. Soon after the sinusoid reverses polarity, the 
arc quenches until the spark-gap breakdown voltage or 
restrike potential is reattained. This repetitive arcing is 
almost always self-sustaining at ac voltage levels of 480 V 
and above. 

Depending on how the fault occurs, it may be de- 
scribed as three-phase, line-to-line, or line-to-neutral. In 
mining, the cable shields and the grounding system of the 
equipment are at the same potential as the system neu- 
tral, and line-to-neutral faults are the most prevalent. 
Line-to-line faults and line-to-neutral faults are unbal- 
anced or unsymmetrical, but the three-phase fault is 
balanced or symmetrical. These three basic fault descrip- 
tions are illustrated in figure 4.32. The impedance, al- 
though very small, is shown to signify its finite value. 



-1 




1 1 — — 

» 

Zq r $ Zg $ Zq 
i — 1 1 1 

1 . — 



Line to neutral Line to line 3 phase 

Figure 4.32.— Basic fault descriptions. 



Fault Analysis 

Fault analysis is a desirable and often mandatory part 
of any mine power-system analysis. As faulting of some 
nature can occur at any time, knowledge of how faults 
affect currents and voltages is necessary to design proper 
protection and to ensure personnel protection. Although 
faults usually occur in mine-system trailing cables, the 
actual fault location and time of occurrence is unpredict- 
able. Consequently, fault analysis is frequently effected on 
a trial-and-error basis, searching for a worst case solution. 
It is necessary to assume a fault location, the configura- 
tion of power-system components prior to the fault, and 
sometimes the system loads. Such an effort can result in 
numerous calculations, to the point where digital comput- 
ers can be extremely advantageous. Nevertheless, the 
results provide invaluable information on which to base 
the design of the mine power system. 

Though not particularly common, fault analysis using 
three-phase faults has distinct advantages. Using this 
method, balanced faults, like balanced loads, can be inves- 
tigated on an equivalent per-phase basis and therefore 
become as simple as faults on single-phase lines. In the 
majority of cases, bolted three-phase faults cause larger 
fault currents than line-to-line or line-to-neutral events. 

Unsymmetrical faults are often of high interest in 
mine power systems. Instances include finding a mini- 
mum fault current or current flowing in the system 
neutral conductors. When an unsymmetrical fault is 
placed on the system, the balance is disrupted. It is 
possible to solve an unbalanced power system by using a 
three-phase diagram with symbols assigned to the quan- 
tities in each phase and carrying the phase solutions 
simultaneously. This complicates the problem enormously, 
but it can be simplified by applying the method of sym- 
metrical components, which reduces the solution of such 
problems to a systematic form. The reduction is particu- 
larly applicable to balanced systems operating under 
unsymmetrical faults. 



SYMMETRICAL COMPONENTS 

The method of symmetrical components provides a 
means for determining the currents and voltages at any 
point of an unbalanced three-phase power system. In this 
method, the unsymmetrical phasors representing the un- 
balanced voltage or current are expressed as the sum of 
three symmetrical phasor sets. These phasor sets or sym- 
metrical components are designated as the positive se- 
quence, negative sequence, and zero sequence. The first 
two consist of three balanced phasors with equal magni- 
tude, set 120° apart. The zero-sequence set has three 
phasors equal in magnitude but operating in the same 
time. The components are illustrated in figure 4.33, where 
the instantaneous values may be determined by projection 
upon the X-axis. The positive-, negative-, and zero- 
sequence components are then employed to solve the 
unbalanced-system problem. These sequences are so com- 
mon in power-system terminology that they are often used 
to describe the quality of system operation. 

It might be asked why the resolution of three phasors 
into nine phasors simplifies the solution of unbalanced 
power systems. The answer is straightforward. The resolu- 
tion results in three symmetrical systems in which each 



99 





V a0 V b0 

) v//- 



v c0 



Figure 4.33.— Positive-sequence, negative-sequence, and 
zero-sequence vector sets. 



balanced phasor set can be treated separately, just as in 
balanced three-phase systems. In other words, the power 
system can be reduced to per-phase values, then analyzed 
separately for each symmetrical component. This analysis 
hinges on the fact that currents and voltages of different 
sequences do not react upon each other: currents of the 
positive sequence produce only positive-sequence voltage 
drops; the same is true for the negative and zero sequences. 
In addition to aiding analysis, the method of symmetri- 
cal components separates electrical parameters into parts 
that can represent better criteria of the controlling factors for 
certain phenomena. For example, the presence of negative- 
sequence current or voltage immediately implies that the 
system is unbalanced, and this can be utilized to detect 
malfunctioning power systems. Grounding phenomena are 
other good criterion examples; neutral current is very closely 
related to zero-sequence components. 



Sequence Components 

The positive sequence for voltage is composed of three 
symmetrical phasors, V al , V bl , and V cl for phases a, b, and 
c, respectively (fig. 4.33). The quantities have equal mag- 
nitude but are displaced by 120° from each other. There- 
fore, following equation 4.26, 



V al = V bl |12po = V cl |240^, 
or rewriting in exponential form, 

» al = »al> 



V - pi 240 V 

v bl - ^ v al> 



V cl = e* 120 V al . 



The unit vector, e* 120 , is used so frequently that it is given 
the symbol "a" (not to be confused with the transformer 
turns ratio), where 



and 



a = e^ 120 = 1 | 120° 

& 2 _ ^120^120 _ eJ240 



(4.36a) 
(4.366) 



Equations 4.35 and 4.37 also relate to the standard practice 
of symmetrical-component calculations; equations are al- 
ways expressed in terms of the phase a quantities. 

There are several mathematical properties of the unit 
vector a that are useful in computations: 

1 = e* = 1.0 + J0.0, 

a = e* 120 = -0.5 + jO.866, 

a 2 = (J 240 = -0.5 - jO.866, 

a 3 = e* 360 = 1.0, 

a 4 = e* 480 = e* 120 = a, 

a 5 = e* 600 = e j240 = a 2 ; 

and for specific calculations: 

1 + a 2 + a = 0, 

a - a 2 = V3e*° = jV3, 

a 2 - a = V3e " j9 ° = -jV3, 

1 - a = 3e - j30 = 1.5 - jO.866, 

1 - a 2 = V^ 30 = 1.5 + jO.866. 

These allow easy conversion to simpler forms when sym- 
metrical components are being manipulated mathemati- 
cally. For the negative and zero sequences (fig. 4.33), the 
symmetrical voltage sequences can be written 



V a2 = V b2 1 -120° = V c2 1 120' 



and 



V„„ = V hn = V„ n . 



(4.38) 



(4.39) 



Rewriting these equations in terms of the unit vector, a, it 
is found that for the negative sequence, 



V^ = V^, 



Thus the positive-sequence vectors (equations 4.35) are 
customarily written as 



(4.40a) 
(4.406) 
(4.40c) 

(4.41a) 
(4.416) 
(4.41c) 



It is important to note that in all three sequence 
systems, the subscripts denote specific components in each 
phase (a, b, or c). Furthermore, the subscripts, 1, 2, and 
signify whether that component is part of the positive-, 
negative-, or zero-sequence set. Using the same reasoning, 
symmetrical-component equations can also be written for 
current. 



Vt.ot; 






V b2 = aV a2 , 








*c2 = a *a2> 


(4.35a) 








(4.356) 


and for the 


zero 


sequence, 


(4.35c) 






* a o = »aO> 
' bO = "aO> 


s given 
iformer 






V c0 = V a0 . 



'al — *al> 

V bl = a 2 V al , 
V cl = aV al . 



(4.37a) 
(4.376) 
(4.37c) 



Sequence-Quantity Combinations 

The total voltage or current of any phase is equal to 
the vectorial sum of the corresponding components in that 



100 



phase. Figure 4.34 illustrates this concept for three unbal- 
anced voltage phasors, V a , V b , and V c . Expressed mathe- 
matically, 

V a = V a0 + V al + V^, (4.42a) 

V b = V b0 + V bl + V b2 , (4.426) 

V c = V c0 + V cl + V c2 . (4.42c) 

Substituting in the equivalent values given by equations 
4.37, 4.40 and 4.41, equations 4.42 become 

V a = V a0 + V al + V^, (4.43a) 

V b = V a0 + a 2 V al + aV^, (4.436) 

V c = V a0 + aV al + a 2 ^. (4.43c) 

These equations state that an unbalanced system can be 
defined in terms of three balanced phasor sets. In other 
words, positive-, negative-, and zero-sequence components 
of phase a can be added together to obtain the unbalanced 
phasors. Following convention, the equations are ex- 
pressed only in phase a quantities. 

Similarly, three unbalanced voltages or currents may 
be resolved into their symmetrical components. Consider 
the zero sequence first. By adding equations 4.43a, 4.436, 
and 4.43c together, 

V a + V b + V c = 3V a0 + (l + a 2 + a)V al + (l + a + a 2 )^. 

Since 1 + a 2 + a = 0, 

V a0 = I(V a + V b + V c ). (4.44a) 

If equation 4.436 is multiplied by a and equation 4.43c by 
a 2 and these results are added to equation 4.42a, 

V a + aV b + a 2 V c = (l + a + a 2 )V a0 + 3V al + (l + a 2 + a)V a2 . 



Therefore, 



V al = |(V a + aV b + a 2 V c ), 



(4.446) 



which relates the positive-sequence component of phase a 
to the unbalanced vectors. Finally, for the negative se- 
quence, if equation 4.436 is multiplied by a 2 and equation 
4.43c by a, 

V a + a 2 V b + aV c = (1 + a 2 + a)V a0 + (1 + a + a 2 )V al 
+ 3V a2 . 

Then, V a2 = |(V a + a 2 V b + aV c ). (4.44c) 

Equations 4.44a, 4.446, and 4.44c are therefore the reverse 
of equations 4.43a, 4.436, and 4.43c; they allow the sym- 
metrical components to be written in terms of the unbal- 
anced phasors. 



Symmetrical-Component Relationships 

Currents in equivalent delta-connected and wye- 
connected loads or sources form a good basis to illustrate 
the existence of symmetrical components in three-phase 
circuits. Consider the two loads shown in figure 4.35, 
where I ab , 1^, and I ca are the three phase currents and I a , 



I b and I c are the line currents. These may all be assumed 
to result from an unbalanced condition. 

At the three terminals of the delta load, the following 
relationships are satisfied by Kirchhoff s current law: 



•'■a = ■'■ab — ^ca> 


(4.45a) 


lb — Ibc - Iab> 


(4.456) 


■*-c A ca -4>c' 


(4.45c) 



The zero-sequence currents of the wye-connected load are 
(using equation 4.44a): 



IaO = 3 da + lb + U 



(4.46) 



Substituting equations 4.45 into equation 4.46 it is found 
that 

IaO = IWca + Iab + Ibc) " (Lb + Ibc + lj\ = 0. (4.47) 

This shows that the zero-sequence current of a three-phase 
circuit feeding into a delta connection is always zero. In 
addition, the currents to a three-phase wye-connected load 
with a floating neutral (fig. 4.35B) can have no zero- 
sequence component. Simply, a neutral-return circuit 
must be available for zero-sequence currents to flow. 
However, zero-sequence current may circulate in a delta 
connection without escaping into a neutral conductor (see 
figure 4.35A, note directions of I ab , 1^, and I ca ). 




Figure 4.34.— Symmetrical component addition to obtain 
unbalanced three-phase set. 





Figure 4.35.— Equivalent delta-connected {A) and wye- 
connected (B) loads. 



101 



For transforming positive-sequence line currents to 
phase currents, it can be shown from 



J ai = 3 da + alb + a 2 I c ). 



(4.48) 



which applies equation 4.446 to current, that 

J 
V3 

Using a similar process for negative-sequence currents, 

Iab2=^Ia2- (4-50) 



Iabi = -75lai- (4-49) 



73 



When the foregoing is applied to line-to-neutral and line- 
to-line voltages for figure 4.33, the transformation equa- 
tions are 



V ab0 = 0, 

Vabi=jV3V al , 

V M = -W3 V a5 



(4.51a) 
(4.516) 
(4.51c) 



where V ab0 , V abl , V ab2 = zero-, positive-, and negative- 
sequence line-to-line voltages, 
V, 
V a0 , V al , V^ = zero-, positive-, and negative- 
sequence line-to-neutral volt- 
ages, V. 

These equations demonstrate another general relation- 
ship of zero-sequence components: a line-to-line voltage, 
however unbalanced, can have no zero-sequence compo- 
nent. Line-to-neutral voltages, on the other hand, may 
have a zero-sequence value. 



Symmetrical-Component Impedance 

Before the solution of unbalanced system problems 
can be discussed, the concepts of impedance under the 
influence of symmetrical components need to be covered. 
Impedance relates the current in a circuit to the impressed 
voltage. Symmetrical-component impedance behaves in a 
similar manner, except that it is sometimes affected by 
additional parameters. There are three likely cases for a 
power system: an unbalanced static network, a balanced 
static network, and the balanced nonstatic network. All 
these impedance values are created by the fact that 
positive- and negative-sequence currents produce only 
positive- and negative-sequence voltage drops, respec- 
tively. The flow of zero-sequence currents in a neutral can 
result in an impedance that is apparently greater than the 
actual impedance. 

In an unbalanced static network, the sequence imped- 
ances in a particular phase are equal, but not necessarily 
equal to those in another phase: 



Z a o - Z al - Z^, 


(4.52a) 


Zbo = Zn,! = Zbg, 


(4.526) 


Zco = Z cl = Z c2 , 


(4.52c) 



where Z a0 , Z b0 , and Z c0 are symmetrical-component imped- 
ances for the zero sequence; Z al , Z bl , and Z cl are 
symmetrical-component impedances for the positive se- 
quence; and Z^, Z b2 , and Z c2 are symmetrical-component 
impedances for the negative sequence. 

The balanced nonstatic network is given by 



Z a o = Z b o = Z c0 , 


(4.53a) 


Z a l = Z b i = Z cl , 


(4.536) 


Z a 2 = Z b2 = Z c2 . 


(4.53c) 



This states that in a balanced nonstatic network the imped- 
ances in a sequence are equal, but not necessarily equal to 
the other sequence component impedances. Cables and pow- 
erlines are included in this case, and specifically, 



Z a l ~ Za2 =£ Z a0 . 



(4.54) 



The last likely case is the balanced static network, 
where 



Z = Z K = Z„ 



(4.55) 



It should be obvious that this is a situation where sym- 
metrical components would not normally be applied. 

As a general rule, positive- and negative-sequence 
impedances of a power system are on the same order of 
magnitude, but the system zero-sequence impedance may 
vary through a very wide range. This range is dependent 
upon the resistance-to-reactance ratio as seen by the 
zero-sequence current. 



Fault Calculations 

One of the most significant uses for the method of 
symmetrical components is the computation of voltages 
and currents resulting from unbalanced faults. The three- 
phase diagram in figure 4.36 represents a simple power 
system with a four-wire wye-connected source. The imped- 
ance of each phase conductor is Z, while Z is the neutral- 
conductor impedance. A bolted line-to-neutral fault is 
occurring on a phase a (an x signifies the fault). The 
resulting current in the fault, If, is of interest, and the 
following shows how symmetrical components can be used 
to find its value. 



V V f ■ / 

-o VWV o X 



Fault 




Figure 4.36.— Three-phase system with line-to-neutral fault. 



102 



It is apparent from figure 4.36 that the line currents 
under the fault condition are 

I a = ^ I b = 0, I c = 0, (4.56) 

where If = current in fault, A, 

and I a , I b , I c = unbalanced line currents, A. 

Applying equation 4.43, the symmetrical components of 
these currents are 



the current is 3I . Zero-sequence impedance, Z a0 , is there- 
fore greater than Z. As was implied in the preceding 
section, the quantification of Z a0 , or just simply Z , is not 
an easy matter. However, in order to limit the amount of 
fault current flowing in mine power-system neutrals, large 
resistances are placed in the neutral circuit. With this in 
mind, the resistance-to-reactance ratio of the neutral im- 
pedance, Z n , is very large, and in this instance for mine 
power systems under line-to-neutral faults, the impedance 
seen by the zero-sequence current can be approximated as 



therefore, 



I al = | (I a + al b + a 2 I c ) = I If, (4.57a) 
ia 2 = |(Ia + a 2 I b + ai c ) = |if, (4.576) 

Ial=ia2 = IaO = 5lf (4-58) 



Z = 



Z + 3Z, 



(4.60) 



To define the fault completely, it must be known 
whether the fault between line a and the neutral is a dead 
short or exhibits an impedance. Although all faults have a 
finite impedance, the faulting assumption states that it is 
bolted. Therefore ,Jault impedance is zero and the voltage 
across the fault, V fn , is also zero. With this, the current 
through the fault, If, can be described. However, to perform 
the required computations, it is necessary to know the 
force driving the fault current and the impedance existing 
between this driving potential and the fault location. 

The source, V, is the driving potential, and it can be 
assumed as purely positive sequence. It is also assumed from 
figure 4.36 that the source has negligible internal imped- 
ance (in practical situations, however, the source impedance 
is of great importance). Therefore, the source line-to-neutral 
potentials are equal to the terminal voltages: 



V =VW.=VwV =V 

'an v a> v bn v b> v cn v c 



where V^V^V^ 



= terminal line-to-neutral voltages 
and V Q , V b , V c = corresponding ideal-source poten- 

tials. 



The impedances involved are simply the line impedance of 
phase a (Z) and the neutral impedance (Z n ). With these 
parameters known, the process is now to convert the 
unbalanced system to symmetrical components, solve the 
problem in terms of these balanced vectors, and then 
reconstruct the result for the fault current. 

Following convention, all work is performed in phase 
a quantities. Notwithstanding, phase a contains the line- 
to-neutral fault; thus, it is the only phase involved. Since 
only positive-sequence voltage is supplied by the source, 
the symmetrical components of the driving potential are 



'anl — *an> *an2> — U, *anO — "■ 



(4.59) 



The impedance to positive-sequence or negative- 
sequence current in any of the three lines is equal; thus, 
for phase a, 

Z a i = z, Z a2 = z. 

Zero-sequence current follows a different path from posi- 
tive or negative sequence. From the source to the fault, 
zero-sequence current, I , exists in each line, but from the 
fault back to the source (through the neutral conductor) 



Loop equations for each sequence current can now be 
expressed for figure 4.36. If a voltage is assumed to exist 
across the fault, for phase a, 



'anl 



= Z al I al + V n , 



v an2 = z a2 I a2 + V^ = 0, 

V a „0 = Z I a0 + V ro = 0, 



(4.61a) 
(4.616) 
(4.61c) 



where V anl , V an2 , V^q = sequence voltages for source, V, 
I al , I a2 , I a0 = sequence components of fault 
current, A, 
Vfi, Vf2, VfQ = sequence components of volt- 
age across fault, V, 
and Z al , Z a2 , Z = sequence impedance seen by 

fault current, fl. 

Equation 4.59 generalizes the fault condition and is prac- 
tical because of fault impedance. However, a bolted fault 
has been assumed to exist; thus, 



v n = V» = V ro = 0. 



(4.62) 



All input to the problem is now available, and simulta- 
neous solution of equations 4.57 through 4.62 shows that 

Van = V anl = § ZIf + § ZIf + | Z If (4.62a) 



or 

but 

then 



If = 



Va (2Z + Z ) ' 



(4.626) 



Z = Z + 3Z r 



If = 



Z + Z„ 



(4.62c) 



Consequently, symmetrical components have been em- 
ployed to solve this unbalanced faulting problem. 

This work can easily be expanded to cover other 
unbalanced faulting problems, and the process can be 
employed to solve any unbalanced three-phase or even 
polyphase condition. Because fault analysis is imperative 
in protective-device sizing, additional discussion can be 
found in chapter 10. 



POWER TERMINOLOGY 

If the sum of the electrical ratings is made for all 
equipment in a power complex, the result will provide a 



103 



total connected load. The measure could be expressed in 
horsepower, but the electrical quantities of kilowatts, 
kilovoltamperes, or amperes are more suitable units. Note 
that the connected horsepower can be converted to con- 
nected kilowatt simply by multiplying by 0.746. Many 
loads operate intermittently, especially mining production 
equipment, and other equipment operates at less than full 
load. Accordingly, the demand upon the power source is 
less than the connected load. This fact is important in the 
design of any mine power system, as the system should be 
designed for what the connected load actually uses, rather 
than the total connected load. Obviously, these consider- 
ations have great impact on power-system investment or 
the capital required to build the system. 

Because of the importance of assessing equipment 
power demands, the Institute of Electrical and Electronics 
Engineers (IEEE) has standard definitions for load combi- 
nations and their ratios. The important ones follow (3). 

• Demand is the electrical load for an entire complex 
or a single piece of equipment averaged over a specified 
time interval. The time or demand interval is generally 15 
min, 30 min, or 1.0 h, and demand is generally expressed 
in kilowatts, kilovoltamperes, and amperes. 

• Peak load is the maximum load consumed or pro- 
duced by one piece or a group of equipment in a stated 
time period. It can be the maximum instantaneous load, 
the maximum average load, or (loosely) the maximum 
connected load over the time period. 

• Maximum demand is largest demand that has 
occurred dur ing a specif ied timejseriod. 

• Demand factor is the ratio of the maximum demand 
to the total connected load. 

• Diversity factor is the ratio of the sum of the 
individual maximum demands for each system part of 
subdivision to the complete system maximum demand. 

• Load factor is the ratio of the average load to the 
peak load, both occurring in the same designated time 
period. This can be implied to be also equal to the ratio of 
actual power consumed to total connected load in the same 
time period. 

• Coincident demand is any demand that occurs si- 
multaneously with any other demand. 

All these definitions may be applied to the units of 
average power, apparent power, or current. Thus they are 
invaluable in power-system design. A few examples are in 
order to illustrate their versatility. 

Consider a feeder cable supplying several mining 
sections in an underground mine. The sum of the con- 
nected loads on the cable, multiplied by the demand factor 
of these loads, yields the maximum demand that the cable 
must carry. When applied to current, this demand would 
be the maximum amperage. Good demand factors for mine 
power systems range from 0.8 to 0.7 depending upon the 
number of operation sections. The lower value is used 



when there are fewer producing units, that is, from two to 
four. The demand factor can be extended to include esti- 
mates of average load. For instance, the sum of the average 
loads on the cable, multiplied by the demand factor, 
provides the average load on the cable. A prime applica- 
tion here is for approximating the current that a conductor 
is expected to carry. If, for example, 10 identical mining 
sections draw 53 A each; the conductors feeding all these 
sections would be expected to carry 

(total average loadXdemand factor) = (average load) 



or 



(10X53 A) (0.8) = 424 A. 



The demand factor and the diversity factor can be applied 
to many other mine electrical areas, such as estimating 
transformer capacities, protective-circuitry continuous 
ratings, and the load that a utility company must supply. 

The load factor can be used to estimate the actual 
loads required by equipment. Here, the total connected 
load multiplied by the load factor is an approximation of 
the actual power consumed. It should be noted that the 
average load factor in underground coal mining tends to be 
rather low, mainly because of the cyclic nature of equip- 
ment operation but also because of the employment of 
high-horsepower motors that are needed to perform spe- 
cific functions but only operate for a small fraction of the 
possible running time. For instance, when cutting and 
loading, a continuous miner will have all motors operat- 
ing, thus have a total connected load of 385 hp or 
(0.746)385 = 287 kW. The average load factor might be 
0.6; therefore, the actual power consumed is (0.6)287 or 
172 kW. The load factor can also be applied to equipment 
combinations. 

The maximum power demand normally forms one 
basis that utility companies use to determine power bills; 
most often, 1 month is the specified time period. Demand 
meters are often installed at the utility company metering 
point. 

Chapter 4 has covered a broad range of fundamentals 
projected towards three-phase power systems in mining. 
Items have included balanced three-phase circuit analysis, 
the per-unit system, the method of symmetrical compo- 
nents, and specific terminology to describe power-system 
operation. Comprehension of this material is vital in order 
to understand many chapters that follow. 



REFERENCES 

1. American National Standards Institute (New York). ANSI 
Standard Device Numbers C37.2. 1978. 

2. Graphical Symbols for Electrical Diagrams Y32.2. 

1970. 

3. Institute of Electrical and Electronics Engineers (New York). 
Recommended Practice for Electric Power Distribution for In- 
dustrial Plants. Stand. 141-1986. 



104 



CHAPTER 5.— BASIC SOLID-STATE DEVICES AND INSTRUMENTATION 



Through the advancement of technology, the motor- 
generator (m-g) sets and Ignitron rectifiers for power conver- 
sion used in early mining have been all but replaced by 
semiconductor devices, except for m-g sets and synchronous 
rotary converters in specific surface mining equipment. 
Equipment employing semiconductors exclusively is often 
termed solid state or static. In mine power systems the use of 
semiconductors has grown from simple rectification (the 
conversion of power to direct current (dc)) to include such 
areas as motor and equipment control, protective relaying, 
and lighting power supplies, not to mention extensive use in 
communications and instrumentation. 

Since the topics of solid-state devices and basic instru- 
mentation are closely related, they are introduced to- 
gether in this chapter. The discussion will be primarily 
informative rather than theoretical. 



SEMICONDUCTORS 

Semiconductors are nonmetallic elements that are 
characterized by relatively poor conductivity. Silicon is the 
most popular and germanium the second most important 
semiconductor in electrical or electronic applications. 
Semiconductors are useful in electrical circuits because 
they can pass current in two different conduction modes 
when impurities or imperfections exist in their crystal 
lattices. The process of carefully adding impurities to a 
pure or intrinsic semiconductor crystal while it is being 
grown is called doping. The impurities are selected for 
their size, so they will fit into the crystal lattice and 
provide either an excess or a deficiency of electrons. 

For example, when a few parts per million of arsenic 
atoms are added to germanium, or antimony atoms are 
added to silicon in the crystal structure, an overabundance 
of free electrons is created. The result is a net negative 
effect, and the crystal is termed an n-type semiconductor. If 
a potential is placed across the impure crystal, conduction 
occurs primarily through an apparent drift of these free 
electrons. On the other hand, if indium or gallium is used 
to dope germanium or silicon, a deficiency of electrons 
exists, and an excess positive charge is created in the 
doped crystal. Thus, it is called a p-type semiconductor. If a 
potential is applied, the atoms conduct current by an 
apparent movement of electron sites or holes. These holes 
are places in the crystal lattice where an electron can be 
held temporarily. When there is an abundance of holes, 
free electrons generated within the crystal can quickly 
recombine with available atoms. 

The free electrons in the n-material and the holes in 
p-material are known as majority carriers. However, be- 
cause of thermal and other energies, free electrons are also 
found in a lesser amount in the p-type and a few holes exist 
in n-type semiconductors. These are called minority carri- 
ers. Nevertheless, even with the excess charge, both semi- 
conductor types are electrically neutral. 



DIODES AND RECTIFIERS 

The operation of most semiconductor devices is depen- 
dent upon a p-n junction, which is the boundary formed 
when a piece of p-type material is joined with a piece of 



n-type. In the actual production, a single semiconductor 
crystal (or monocrystalline material) is grown so that part 
is doped to create a p-type region, with the balance doped 
for n-type. A solid-state diode or rectifier has one p-n 
junction; it is a device that readily passes current in one 
direction but does not permit appreciable current in the 
opposite direction. The symbol for a diode or rectifier is 
given in figure 5.1 A. 

Figure 5.15 is a simple model of a diode that can be 
used to explain p-n junction electrical operation. When 
the two semiconductor materials are joined, a charge 
redistribution occurs. Both the p-region and the n-region 
contain a high concentration of majority carriers. Elec- 
trons from the n-material diffuse across the junction to the 
p-material; similarly, holes migrate from the p-material 
into the n-material. The net result of this diffusion is a 
depletion region with negatively charged (acceptor) ions 
on the p-side and positively charged (donor) ions on the 
n-side of the junction. The electric field across the deple- 
tion region is established and opposes further majority- 
carrier diffusion, but the field creates a minority-carrier 
flow across the junction in the opposite direction. 

Current caused by majority-carrier diffusion is called 
injection current, I ls and that from minority carriers, 
saturation current, I s . If no external voltage is applied to 
the p-n device (fig. 5.2A), the junction is in equilibrium 
because the net hole and electron flow across the junction 
is zero. In other words, injection current equals saturation 
current. However, if an external voltage is applied with a 
polarity such that the p-region is positive with respect to 
the n-region (fig. 5.25), the depletion-region electric field 
is decreased, and a large number of majority carriers are 
able to cross the junction and diffuse toward the device 
terminals. Hence, injection current is substantially in- 
creased, and because saturation current remains constant, 
the result is current flow in the external circuit. In this 
case, the external voltage polarity is called forward bias, 



p-type June 


tion n-type 


L±J Ld 


L+J Ld 


69 eeeee 


© ® © © ® © © 


++++++ 


— — — — — — 


0990969 


©©©©©©© 


+ + + ++ + 





Depletion region 

Figure 5.1. —Symbol (A) and operation (8) of a p-n junction 
device. 



Js 
-M- 






Js 



I s =l! Is <:[ I I S >I I 

A No external voltage B Forward bias C Reverse bias 

Figure 5.2.— Bias conditions and current flow for a diode. 



105 



and the current is forward current. Conversely, reverse 
bias, an applied voltage of reverse polarity (fig. 5.2C), 
opposes majority-carrier diffusion by enforcing the 
depletion-region electric field, and current is greatly re- 
duced. As saturation current is still constant, the external 
reverse current is primarily a result of minority-carrier 
diffusion and is therefore very small. 

Because there are many more majority carriers than 
minority carriers, the injection current, under forward 
bias, is orders-of-magnitude greater than the constant 
saturation current. As external circuit current is the 
algebraic sum of injection and saturation currents, for- 
ward current is significantly greater than reverse current. 
Furthermore, to enhance this one-directional current phe- 
nomenon, junctions are manufactured in which one side of 
the junction is more heavily doped than the other. Forward 
current is then mainly a result of majority carriers from 
the more heavily doped region. 

The arrow portion of the diode symbol (fig. 5.1A) 
points in the same direction as forward current. As a 
carryover from vacuum-tube terminology, the side symbol- 
ized by the arrow is also called an anode (the p-region), 
with the opposite terminal, the cathode (the n-region). 



the external current is about equal to the saturation 
current. Therefore, by placing a reverse bias across the 
device and measuring the resulting reverse current, the 
forward current can be predicted. 

The foregoing equations result in the theoretical 
curve, termed a characteristic curve, which is given in 
figure 5.3. This curve diverges from that for an actual 
diode in one main aspect, the breakdown of the p-n 
junction noted at point c. Here, the external voltage meets 
the limit capabilities of the junction, and a greater reverse 
voltage will create an avalanche current that can destroy 
the device. As a result, p-n junctions normally require a 
rating for maximum reverse voltage or peak inverse volt- 
age (PIV). Zener diodes are of special interest as they 
operate in this avalanche current area to regulate an 
applied dc voltage. 

As long as the p-n junction is operated within the 
limits of its reverse voltage and forward current, the device 
can be represented by a very low resistance for forward- 
bias conditions and a high resistance during reverse bias. 
Ideally, and for the majority of applications, a diode can be 
assumed to have zero resistance under forward bias and 
infinite resistance under reverse bias. 



Diode Equations 



Rectifier Circuits 



The number of minority carriers is dependent upon 
temperature and the difference in energy levels between 
the p- and n-regions. If the energy difference is constant, 
the concentration of minority carriers plus the saturation 
and reverse currents will vary exponentially with temper- 
ature. Therefore temperature is a limiting factor in diode 
operation, and the maximum rated current of a given 
device is determined by the heat-dissipating properties of 
the device mounting system. 

The formula relating external and saturation current 
with the energy difference and temperature is 



I = -I s (e KT - 1), 



(5.1) 



where I s = saturation current, A, 

q = charge of one electron, 16 x 10 ~ 20 C, 
V = voltage across junction Qess than external volt- 
age, but approximately equal to it), V, 
qV = energy difference between p- and n-materials, 
K = Boltzmann constant, 1.38 x 10 ~ 23 J/K, 
T = junction temperature, K. 



and 



At room temperature (300 K), 



I = -I s (e 39V - 1), 



(5.2) 



or at other temperatures, 



I = -I s (e (39VT ^ - 1), (5.3) 

where T 1 = 300 K, 

and T 2 = other temperature, K. 

The negative sign for the saturation current denotes it as 
flowing in the opposite direction to forward current. The 
equations relate that if voltage is 0.1 V or more negative, 



A rectifier can be considered as a diode specifically 
designed or applied to convert power to dc. The principal 
application in mining is to use the unilateral properties of 
tbe rectifier for direct alternating current (ac) power 
conversion. With single-phase ac, there are three basic 
rectifier circuits to perform this function: half-wave, full 
wave, and bridge. 

Figure 5.4A illustrates the circuit of a simple half- 
wave rectifier in which a transformer magnetically cou- 
ples the source to the rectifier. This could also be direct, 
unisolated source connection. With a sinusoidal voltage 
input (fig. 5.4B), the rectifier acts as a switch. When 
forward biased (positive anode with respect to the cath- 
ode), the load, R, is electrically connected to the source, 
but during reverse biasing it is disconnected. In other 
words, low and high resistances to current exist with 
respect to the bias condition. These resistances create a 
pulsating dc waveform across the load, as shown in figure 
5.4C. This variation of voltage is often termed ripple. 



~]~ Theoretical 
,'N 
/ Actual 




/ b 
1 c 

i 
i 
i 
i 
i d 



< VOLTAGE, ICf'V 



u 
cr_ 

=> 
v 



Figure 5.3.— Diode or rectifier characteristic curve. 



106 



Only the positive portions of the input sinusoid ap- 
pear in the pulsating dc output, and as a result, the 
conversion efficiency of the half- wave rectifier leaves much 
to be desired. The single-way full-wave rectifier is a 
method of rectifying both the positive and negative por- 
tions of a sinusoidal voltage input, and it can be analyzed 
as two half-wave rectifiers. The circuit shown in figure 
5.5A utilizes a center-tapped transformer secondary. 
When referenced to ground, the V 2 and V 2 ' waveforms (fig. 
5.5J3) are then 180° out of phase. Therefore, one rectifier is 
conducting current (forward biased) while the other is not 
(reverse biased). The consequence is pulsating dc power to 
the load during both the negative and positive portions of 
the ac input (fig. 5.5C). Conversion efficiency is greatly 
improved over half-wave circuits. 

Full-wave rectification can also be obtained with the 
bridge rectifier. As shown in figure 5.6A, the circuit 
employs a transformer with a single secondary and four 
rectifiers. During either the positive or negative portions 
of the input waveforms, two of the rectifiers are effectively 
in series with the load resistance. For instance, when the 
top secondary transformer rectifiers D 2 and D 3 are forward 
biased but D x and D 4 are reverse biased, current flows 
from the top seco ndary ter minal thro ugh D 2 , R L , and D 3 
back to the transformer. The rectifier biasing condition 
reverses with the transformer secondary polarity (figure 
5.6B, bottom), but the current through the load has the 
same direction. Hence, the same full-wave pulsating dc 
waveform in figure 5.6C appears across the load with only 
half of the secondary turns needed for the single-way 
full-wave rectifier. 

Although the output of these three basic rectifier 
circuits is effectively dc and the current flow is in only one 
direction, the voltage fluctuation or ripple is often too 
great to be useful. Consequently, filtering is required to 
change the pulsating voltage to a relatively ripple-free 
potential. This filtering action is provided by inductors in 
series with the load, or capacitors shunting (in parallel 
with) the load, or both. Each of these methods will smooth 
the voltage output. An example of this filtering is shown 
in figure 5.7. It will be shown later that such filtering is 
not needed for dc mining equipment. 



I W T ° 

c t L ♦ 



out 



^d- 



A r\ AT v, 



out 



A B C 

Figure 5.4.— Half-wave rectifier circuit and waveforms. 




V 2 J 



v out V2 I 






/YY^ 



A B C 

Figure 5.5.— Single-way full-wave rectifier waveforms. 



zf\f 




^V 




rYY\ lv, 



'out 



Figure 5.6.— Bridge rectifier circuit and waveforms. 



Cooling 



It was stated earlier that the operation of a p-n 
junction is highly dependent upon temperature. It follows 
that there exists a maximum temperature beyond which 
the device will be destroyed if operated. Such a point is 
called the maximum junction operating temperature. For 
silicon semiconductors, this temperature is usually 
around 175° to 200° C, for germanium, 85° to 110° C, but 
the maximum varies according to the individual device 
and manufacturer. 

The temperature at which the junction operates is 
dependent upon the power dissipated in the junction, the 
ambient temperature, and the ability of the device to 
transfer heat to the surrounding environment. Devices 
designed and operated for small currents usually do not 
need cooling assistance. However, adequate external cool- 
ing is required in p-n junctions dissipating 1 or more 
watts. The simplest method is to mount the semiconductor 
case securely on a heat sink, which is commonly metal 
with a large surface area. Thermally conductive washers, 
silicon compounds, and correct bolting pressure allow good 
heat transfer from the device to the heat sink, and air 



Series 
Rectifier ^ inductor 
output \ \ 



Capacitor 
shunting load 




Figure 5.7.— Example of filtering a rectifier output. 



convection transfers heat to the surrounding atmosphere. 
In high-power applications, forced-air cooling of the heat 
sink is sometimes employed to increase heat dissipation 
further. 

Figure 5.8 illustrates a rectifier using a heat sink for 
this purpose. The diagram in figure 5.9 represents the 
typical relationships in all solid-state devices between the 



107 




Heat sink 
Figure 5.8.— Heat sink cooling. 



Collector 

junction 

temperature 

_ Case 
temperature, T c 

_ Heat sink 
temperature, T s 

Ambient 
temperature, T a 




JL*. 



Absolute-zero 



temperature 
Figure 5.9.— Heat sink thermal relationships. 



solid-state device, its heat sink, and the surrounding envi- 
ronment. The following equation relates these parameters: 



T j = 0jaPd + T a» 



(5.4) 



where Tj = junction temperature, °C, 

T a = ambient temperature, °C, 

P d = power dissipated by junction, W, 
and ja = ambient-to-junction "thermal resistance," °C/W. 

The last item, thermal resistance, is actually composed of 
three parts, as shown in figure 5.9, 

0ja = 0jc + Acs + C (5-5) 

where jc = junction-to-case thermal resistance, °C/W, 

flea = case-to-heat-sink thermal resistance, °C/W, 
and flga = heat-sink-to-ambient thermal resistance, 
°C/W. 

The junction-to-case and the heat-sink-to-ambient thermal 
resistances are almost always available from manufactur- 
ers. The thermal resistance between the device case and 
the heat sink can be neglected if the mounting is carried 
out correctly as described here. 

Junction power can be found by the relationship 



*d — Imax » f» 

where I,,,^ = maximum forward current, A, 
and V f = junction forward voltage drop, V. 



(5.6) 



The junction forward voltages normally range from 0.5 to 
0.75 V for silicon and from 0.2 to 0.3 V for germanium, but 
typical values for specific devices are also available from 
manufacturers. When the total thermal resistance, ja , is 
known, the operating junction temperature can be calcu- 
lated and compared with the maximum allowed. 

Overloads 

The thermal relationship of figure 5.9 shows three 
capacitances, Cj, C c , and C s , which are the thermal capac- 
itances of the p-n junction, the device case, and the heat 
sink, respectively. Thermal capacitance resists changes in 
temperature in the same way that capacitance restricts 
voltage change. For the p-n junction, Cj is usually very 
small; hence, its time constant is also small. This means 
that the semiconductor must not be overloaded (excessive 
power dissipation) for more than a few milliseconds; other- 
wise, the device will be destroyed. For this reason, high- 
speed overload protection must be applied to semiconduc- 
tor devices. For rectifiers, the protection takes two forms: 
against excessive overloads and short circuits in load 
currents, and against failure in the rectifier itself (over- 
temperature or excessive voltage). 



THREE-PHASE RECTIFICATION 

Large amounts of dc power at either 250 or 500 Vdc 
are required for locomotives and face equipment in many 
mining operations. When more than a fractional kilowatt 
of dc power is needed from an ac source, a polyphase 
rectifier circuit is employed. The direct voltage is derived 
from three-phase ac power, most often from distribution 
voltages. 

There are specific advantages to using polyphase 
rectifier circuits for dc power. As the number of ac phases 
driving the rectifier is increased (say, above single-phase 
ac), the frequency of output ripple is increased, the inter- 
val between rectifier conduction is decreased, and the 
ripple magnitude in the dc voltage and current waves 
decreases. 

Transformers are almost always used between the ac 
source and the rectifiers. The rectifier transformer per- 
forms one or more of the following functions: 

• To transform the available ac supply voltage to a 
value needed for the desired dc voltage; 

• To provide the number of phases required to obtain 
the desired waveshapes of dc voltage, dc current, and ac 
supply current; 

• To isolate the dc circuit from the ac source; and 

• To limit, through transformer impedance, damag- 
ing overcurrents that might flow during malfunctions. 

It is important to note that the decrease in the rectifier- 
conduction interval also increases the required trans- 
former rating. The transformer utilization factor can be 
defined as the ratio of dc power delivered to the required 
transformer secondary voltampere rating. The utilization 
factor has been found to have a maximum value of 0.520 
when three-phase ac input is used. This implies that from 
a transformer utilization standpoint, the most economic 
rectifier-conduction angle is 120°. 

When power rectifiers are mentioned today, the refer- 
ence is almost invariably to solid-state units using silicon 



108 



rectifiers as the rectifying elements. Indeed, the silicon 
rectifier is virtually the only type considered for mine 
power installations. While there are many possible recti- 
fier circuits, only two or three types are found in mining 
equipment. Circuits for silicon rectifiers are selected to 
make the most efficient use of the transformer, and the 
results usually are the single-phase full-wave bridge or the 
three-phase full-wave bridge. The next section will discuss 
three fundamental three-phase rectifier circuits, and it 
will be apparent why the full-wave bridge is popular. 



Phase Phase 
A C 

t 

Phase 
B 




Rectifier Circuits 

Rectifier circuits can be classified as single way or 
double way. The phase currents of the transformer secondary 
(also termed the dc winding) are unidirectional in a single- 
way circuit but alternating in the double-way circuit. 

The simplest three-phase rectifier circuit is the three- 
phase half- wave shown in figure 5.10A, where a delta-wye 
transformer is used, with each leg connected to a rectifier 
anode. The three rectifier cathodes are tied together to 
form the positive dc bus. The neutral point of the trans- 
former winding serves as the negative connection for the 
load, in this case resistance, R. Being a single-way recti- 
fier, each leg of the transformer secondary conducts cur- 
rent unidirectionally. If the load is pure resistance, the 
relationship of output voltage (that across the load) versus 
time is as shown in figure 5.105. Each rectifier conducts 
for the cycle portion in which its anode has a higher 
positive value than the anodes of the other rectifiers. 
Therefore, each rectifier passes current for 120° of the 
input three-phase cycle. Since the current through the 
load is directly proportional to the output voltage, the load 
current has the same waveform as voltage. 

Inspection of the three-phase half- wave output voltage 
shows that the ripple voltage is much lower than the 
single-phase full-wave rectifier circuit. Actually, the rms 
value of the ripple voltage waveform is only 18% of the 
average output voltage (this average voltage is the average 
dc load voltage, V dc ). If rectifier losses are ignored, since 
they are very small for silicon diodes, the dc output voltage 
and the transformer secondary voltage are related by 



A 


B 


C 


A 


B 


V 




V 
i 1 


i 





180° 
out, deg 



360° 



Figure 5.10.— Three-phase half-wave rectifier circuit (A) and 
output voltage waveform (6). 



Phase Phase 
A C 



D, D 3 D 5 




D 2 D 4 D 6 




cut, deg 



V dc = 0.827V max = 1.17V™, 



(5.7) 



where V dc = average dc output voltage, V, 

V max = peak value of voltage applied to rectifier 
circuit, V, 

and Vj^g = rms value of voltage applied to rectifier cir- 
cuit, V. 

Both V max and V^g are line-to-neutral voltages. Note that 
the fundamental frequency is three times the ac line 
frequency. As a result, any filter components required to 
lower the ripple voltage further can be much smaller than 
in single-phase rectifiers. 

The relationships presented here for the three-phase 
half-wave rectifier apply only to ideal transformers and 
rectifiers. In actual circuits, the voltage drop caused by dc 
current and the transformer secondary-winding resistance 
creates a dc component that pushes transformer magnetic 
operation toward saturation. Consequently, this simple 
three-phase rectifier circuit is seldom used. 

Output ripple can be further reduced by a three-phase 
full-wave rectifier circuit, connected as shown in figure 
5. 11 A. This circuit is also called the three-phase bridge or 




180° 
C cut, deg 

Figure 5.11.— Three-phase full-wave rectifier circuit (A) with 
input (0) and output (C) voltage waveforms. 



a six-phase rectifier. Being a two-way rectifier, the mag- 
netic saturation problem in the transformer is not present. 
Furthermore, this configuration retains the advantage of 
120° conduction for transformer economy, plus a funda- 
mental ripple frequency of six times the ac source fre- 
quency. These characteristics make this double-way recti- 
fier circuit of great practical value, and it is the most 
popular configuration for dc power in mining. The trans- 
former dc winding may be either wye or delta connected. 
In the full-wave rectifier circuit, each terminal of the 
transformer secondary is connected to two diodes, one at 



109 



the anode and the other at the cathode. The cathodes of 
three rectifiers are common and form a positive dc voltage 
bus, while the common anode connection of the other three 
rectifiers represents the negative dc voltage bus. The load 
is connected between these two common points. 

Each rectifier conducts for 120° of one input cycle, and 
current alternates in each transformer winding. However, 
current flows through a specific combination of rectifiers 
for only 60° of the input cycle. This combination could be 
D x and D 4 with transformer secondary terminals A and B. 
Therefore, the peak-to-peak voltage across the load resis- 
tance appears as six-phase ripple as shown in figure 5.11C. 

Analysis of figure 5.11C shows that the rms value of 
the fundamental component of the ripple voltage is now 
only 4.2% of the average dc output voltage. In addition, the 
average dc output voltage for ideal rectifiers is 



V„„ = 0.95V„„ = 1.34V. 



dc 



(5.8) 



where V dc = average dc output voltage, V, 

V max = peak line-to-line voltage applied to rectifi- 
ers, V, 

and Vrmg = rms value of line-to-line voltage applied to 
rectifiers, V. 

The foregoing circuits are typical of most polyphase 
rectifier circuits, but many additional configurations are 
available. Because mining almost always employs full- 
wave rectifier circuits, coverage of more circuits is beyond 
the scope of this text, but the bibliography can be con- 
sulted if desired. 

Parallel Rectifier Operation 

The current requirements of a rectifier circuit are 
often too large to be handled by one rectifier for each 
circuit element. Two or more rectifiers must then be 
connected in parallel. Direct operation of two silicon 
rectifiers in parallel is very difficult, because unbalance 
between the parallel paths can be caused readily by 
unequal rectifier characteristics (mainly the forward volt- 
age) and by unequal impedance in the bus bars or cables. 
The result is that the rectifier with the least forward 
voltage can be destroyed by overcurrent. To eliminate this 
problem, the parallel rectifiers must be forced into sharing 
the current equally. 

The method used almost exclusively in mining equip- 
ment to force current-sharing employs paralleling reac- 
tors, sometimes called current-balancing transformers. 
Figure 5.12 shows how several rectifiers can be paralleled 
using these reactors. The combination acts as one rectify- 
ing element in a rectifier circuit such as in figure 5. 11 A. In 
figure 5.12, each reactor is a laminated magnetic core 
linked in opposing polarity by the anode currents of two 
rectifiers, and the cores are designed not to saturate at the 
highest expected current. If the two rectifier currents 
become unequal, the current difference excites a magnetic 
flux that induces an aiding voltage. This voltage is in- 
duced in the rectifier leads in a direction that will equalize 
the currents. 



TRANSISTORS 

The principal tool of the electronics industry is the 
amplifier, a device that can increase the power level of an 



input waveform or signal. An amplifier is actually an 
energy converter in which energy from a power supply is 
converted by the amplifier to signal energy. The most 
common device used in amplifiers is the transistor. 

A bipolar transistor is formed in a manner similar to 
that of the junction diode, but it consists of two junctions 
in close proximity and parallel to each other in the same 
crystal. When a p-region is sandwiched between two 
n-regions, the device is termed an n-p-n transistor, the 
model and symbol of which are given in figure 5.13A. 
Similarly, if a thin portion of n-material is bounded by two 
p-regions, the transistor is termed p-n-p, as shown in 
figure 5.14A. As illustrated, each semiconductor region is 
given a name: emitter, base, and collector. 

Transistor Operation 

The operation of the transistor is dependent upon the 
bias voltages across the junctions. If voltages are applied 
to an n-p-n device as shown in figure 5.13B, the emitter- 
base junction is forward biased, and the collector-base 



Reactor 



ac 

input 




dc 
output 



Figure 5.12.— Parallel operation of rectifiers using paral- 
leling reactors. 



Base 



Symbol e C 




Emitter 



Collector 



B V CI 



A B 

Figure 5.13.— An n-p-n junction transistor. 



Base 




Emitter 



Collector 



E C 


Veb 


B V CB 


LVhI 


t±|.lpJ 



A B 

Figure 5.14.— A p-n-p junction transistor. 



110 



junction is reverse biased. These are the normal bias 
conditions. Electrons will flow into the base region, caus- 
ing an excess of majority carriers there. Because the base 
region is thin and the potential existing across the two 
n-regions is much higher than the base-to-emitter poten- 
tial, most electrons from the emitter region diffuse across 
the base and are accelerated into the collector region. The 
electrons drift across the collector and cause current flow 
in the collector circuit. However, a small percentage (typ- 
ically, 5% or less) flows out from the base connection 
because of recombination with holes in the base region. 
This process can be considered amplification since the 
small base current controls the much larger collector 
current. A p-n-p transistor operates on the same princi- 
ple, but here it is hole flow rather than electrons that 
causes the amplification. Consequently, the bias condi- 
tions are reversed from that for an n-p-n (the normal 
conditions are shown in figure 5.14B). 

From the preceding discussion, it would appear that 
either end of the transistor could be called an emitter 
because either hole flow or electron flow creates the 
current amplification, but this is generally not the case. 
Heat dissipation is much larger in the collector-base 
junction because of the greater difference in potential. 
Therefore both p-n-p and n-p-n transistors are designed 
so this heat can be diffused through the collector region. 

As might be expected, a saturation current resulting 
from thermally-generated minority carriers flows across 
the reverse-biased collector-base junction. In the diode, the 
current is designated "I s ;" in a transistor, it is termed 
"Icbo" I n the same manner as for diodes, the increase of 
saturation current with temperature sets the maximum 
operating temperature for a transistor. Heat sinks are 
commonly used in high-power transistor applications to 
diffuse collector-base junction heat and maintain temper- 
ature below critical levels. The same calculations that 
were presented in the preceding section on rectifiers can 
also be applied to transistors to determine a safe operating 
temperature. 

The fraction of constant emitter current that reaches 
the collector is called alpha, a, and the collector circuit 
itself can be considered to be the output circuit. Since as 
much emitter current as possible should be collected, 
alpha should be as close to 1 as possible. When combined 
with Icbo' th e collector current, i c , can be expressed in 
terms of emitter current, i E , as 



in = ai, 



+ Ir 



(5.9) 



Figure 5.15 shows the relationship of these currents. 
However, in practical applications, I CB o * s often so small 
that it can be neglected. 

Since base current controls collector current, an im- 
portant expression can be obtained from figure 5.15 using 
Kirchhoff s current law on either transistor: 



1r = lp - i 



c> 

or i B = i E - (ai E + I CBO ) 

= (1 - a)i E - Icbo- 

In terms of collector current, it can be shown that 

Icbo 



(5.10) 



Ir- = 



1 - a 



Ir + 



1 - 



(5.11) 



The term, a/(l - a), is called beta, /3, and also the dc 
current amplification factor, and 



i c = /3i B + (1 + 0)1 



CBO- 



(5.12) 



This last equation shows the significant effect of temper- 
ature on transistor operation; that is, the temperature- 
sensitive I CBO is multiplied by (1 + /3). Even though a is 
less than 1, ft may range from 20 to 200 for amplifying 
transistors. 

Bipolar-Transistor Amplifiers 

Bipolar transistors can be operated with any one of 
the terminals common to the input and output, thus there 
are three basic circuit arrangements: common-base, 
common-emitter, and common-collector. The most popular 
is common-emitter. 

Illustrated in figure 5.16, the common-base or 
grounded-base configuration employs the emitter and base 
terminals as input, with the collector and base terminals 
supplying output. Current gain, which is the ratio of 
output to input, is usually just less than 1. Because the 
emitter-base junction is forward biased, the circuit has low 
input impedance as viewed from the input terminals. 
Because the collector-base junction is reverse biased, the 
output impedance is high in comparison to the input. 
Hence, voltage and power amplification can be realized. 

Two different circuits, signal and bias, are necessary 
for the operation of either of the two common-base ampli- 
fiers shown in figure 5.1. The bias voltage sources, often 
termed the amplifier power supply, fix the dc level for 
proper operation of the two junctions. If the signal input 
and output are not separated electrically from the bias 
source, as seen in figure 5.16A, the circuit is called a dc 
amplifier. Although it is beneficial in applications such as 
amplifying dc voltages for instrumentation, a signal with 






E C 



'C 



E C 



L CB0 




'B 



L CB0 



A B 

Figure 5.15.— Current relationships for p-n-p {A) and n-p-n (B) 
devices. 



^* p-n-p 





Ri 'i(D IR 



'BB =L V CC 

A B 

Figure 5.16.— Common-base amplifiers. 



cc =b- 



Ill 



dc content or offset can interfere with correct transistor 
biasing. Figure 5.16B illustrates a popular method of 
removing this problem: the use of capacitors to isolate the 
amplifier. The capacitors exhibit high impedance to dc but 
low impedance to ac signals, thus they block input and 
output dc. As the circuit now reacts only to ac signals, it is 
called an ac amplifier. It can be noted that transformers 
can perform a similar function. 

With either the dc amplifier or the ac amplifier, a 
small change in input voltage causes significant variation 
in the injection current across the emitter-base junction. 
As previously discussed, most majority carriers diffuse to 
the collector, causing collector current, i c . If the load 
resistance, R L , is small with respect to the transistor 
output impedance, i c is approximately equal to i E . The 
collector current creates voltage variations across the load 
resistance that can be much larger than the input voltage. 

In the common-emitter transistor arrangement, the 
source signal only supplies current to the base. Because 
base current is much smaller than either the emitter or 
collector current, current amplification or gain, Gj, is 
high. Neglecting Icbo in equation 5.12, the gain is approx- 
imately equal to 



G _ L c _ & { B _ 






(5.13) 



which can be from 10 to several hundred. The input 
impedance is also higher than in common-base amplifiers. 

Figure 5.17 shows a simple common-emitter ampli- 
fier. The control action of the base current can be demon- 
strated by assuming that the base-emitter forward bias is 
increased. This increase creates a corresponding increase 
in emitter-base junction current; thus, collector current is 
raised substantially. Because the base current is approxi- 
mately proportional to but usually much less than collec- 
tor current, base current is the controlling parameter of 
the amplifier. 

The concept of characteristic curves has already been 
introduced in figure 5.3 in the section on diodes and 
rectifiers. Characteristic curves are an extremely useful 
tool for the graphical design and analysis of transistor 
circuits. Four independent transistor parameters control 



the number of necessary curves. When figure 5.17 is used, 
these parameters are as follows: 

• a and /3 increase with V CE , the collector-to-emitter 
voltage. 

• i B is dependent on i c and V CE . 

• i B is not a linear function of i c . 

• When V CE is zero, i c is approximately zero, regard- 
less of i B . 

Consequently, two sets or families of curves are needed: 

1. Collector or output characteristics, i c versus v CE for 
varying values of i B , and 

2. Common-emitter input characteristics, V BE versus 
i B for varying values of V CE . 

Figures 5.18A and 5.18B show typical output and input 
characteristics for an n-p-n transistor connected for 
common-emitter operation. The nonlinear and propor- 
tional properties of the four independent transistor param- 
eters are evident in the graphs. These curves can be 
employed for design and analysis purposes. The analysis 
often uses a load line (the straight line in figure 5.18A) to 
observe dynamic variations of voltage and current. 

The dashed line in figure 5.18A is very important as it 
delineates the safe operation boundary. Manufacturers 
specify maximum permissible collector voltage, current, 
and power dissipation, since outside this area damage to 
the transistor will probably result. As noted earlier, allow- 
able power dissipation must be reduced as temperature is 
increased. 



'B 




t v CE H(-s 

'B> 

h| t—M' 

Figure 5.17.— Common-emitter amplifier. 



J ^W | V CE X 1 ^ 

t0 ySMfi E K U 

I ll M III. 1 



4 



s Safe operation 
boundary 




10 20 

COLLECTOR-TO-EMITTER 
VOLTAGE (V CE ), V 

A Output 





Dynamic 


input 


20 V 


QL 


characteristics 


^-15 V 


LU> 0.5 






>^10V 


t w 4 






^.^5V 


LU> 








1 w 3 








*-<2 








1 IS .2 

<> 1 


r i i 


i 


v CE =o 



30 60 90 120 150 
BASE CURRENT 
(I B ). M 

B Input 



Figure 5.18.— Common-emitter characteristic curves. 



112 



Figure 5.17 illustrates an amplifier circuit with two 
batteries supplying dc for transistor bias, but single dc 
source for all bias voltages is more desirable in practical 
applications. Three bias techniques are frequently used 
for common-emitter amplifiers, and these are shown in 
figure 5.19. Each circuit uses resistors to supply dc bias to 
the base for a center bias condition about which the 
transistor operates. The center condition is termed the 
quiescent point of the amplifier. Of the circuits illus- 
trated, the stabilized bias circuit (O gives the best thermal 
stability, maintaining the quiescent point within a desired 
or specified range regardless of the normal operating 
temperature. The bypass capacitor, shown across the emit- 
ter resistor of the stabilized bias circuit, establishes a 
constant base bias bypassing or acting as a low impedance 
to time-varying voltages. 

The two preceding amplifier configurations employed 
the collector circuit for output. In the common-collector 
arrangement, the output is obtained across a load resis- 
tance in the emitter circuit, as illustrated in figure 5.20. 
Because the source and output voltages are now in series 
but have opposing polarities, the circuit gives high input 
impedance and approximately unity voltage gain, yet 
current gain is about the same as in common-emitter 
amplifiers. A main advantage of the common-collector is 
that the output impedance is about equal to the load 
resistance, which is lower than the preceding two connec- 
tions. This allows the circuit to be adjusted to fit the 
output needs precisely; hence, this circuit can be used for 
impedance matching the output of a source signal to the 
input of another amplifier. 

Field-Effect Transistors 

The n-p-n and p-n-p junction transistors just covered 
contained two junctions. Field-effect transistors (FET's) 
have effectively only one junction but still can operate as 
amplifiers. These devices are voltage controlled, whereas 
bipolar transistors can be considered as current-controlled 
devices. There are two general classifications: junction 
FET's and metal oxide semiconductor FET's. Both have 
very high input impedances, much higher than bipolar 
transistors and approaching the input impedance of vac- 
uum tubes. 

To demonstrate the amplifying action available with 
FET's, consider the cross-sectional model of an n-channel 
junction FET, illustrated in figure 5.21A. The gate-to- 
channel junction is reverse biased by placing the voltage 
V GS between the gate and source terminals as shown. The 
level of V GS establishes a specific size of depletion region 
about the gate semiconductor and within the channel. 
Changing this reverse bias increases or decreases the size 
of the depletion region and decreases or increases the 
available conduction area remaining in the channel. 
Therefore, voltage changes between the gate and source 
terminals can control the allowable current through the 
channel from the drain to the source terminals. The action 
can be employed to amplify voltages or currents. 

The conduction channel in the junction can be either 
n-type or p-type semiconductor, with the gate being p- or 
n-material, respectively. Figures 5.21B and 5.21C give the 
symbols for either junction FET type. An important ad- 
vantage of FET's over junction transistors is that the 
source-to-drain channel is resistive without a diode effect. 
In essence, this allows FET's to be operated as electrically 
controlled resistors. 




r-AM/V- 



neW 



,JL 



A Fixed bias 



B Self-bias 




Bypass capacitor 

• » — o * 



C Stabilized bias 

Figure 5.19.— Bias techniques for common-emitter 
amplifiers. 



oV r 




Figure 5.20.— Common-collector amplifier arrangement. 



p-semiconductor 

gate q .Jj3 



^ gA d 




Vest 



DS 



B n-channel symbol 
Ig 



n-channel 
semiconductor 

A Simple bar model 



V SG { 



D 



SD 



C p-channel symbol 



Figure 5.21.— Model and symbols for junction FET's. 



113 



As an application example, figure 5.22 shows a junc- 
tion FET used in a typical amplifier circuit. The input 
signal is applied across the gate to the source, with output 
taken from drain to source. R s is employed to set the 
proper dc quiescent point bias for the gate, and the 
capacitor in the source circuit bypasses ac, thus maintain- 
ing the bias level. 

In metal oxide semiconductor FET's (or MOS-FET's), 
the depletion region used in the junction FET is replaced 
by a thick layer of silicon oxide, a good insulator, and the 
semiconductor employed for the gate is replaced by a metal 
conductor, thus forming a high-quality capacitor. A model 
of a MOS-FET, including the symbols, is given in figure 
5.23. The operation of these transistors is similar to that of 
junction FET's but much more complex. 

The preceding information on transistors is intended 
as just an introduction to a few important devices. For 
complete information, the bibliography must be consulted. 
The coverage here is justified because transistors are an 
extremely important, but often hidden, segment of mine 
power systems. The next section will cover another device 
that has revolutionized the control of electrical machinery. 



SILICON-CONTROLLED RECTIFIERS 

In past few years, the use of solid-state power equip- 
ment in mining has accelerated. One primary reason has 
been the introduction and acceptance of static or solid- 
state starting of conveyor-belt drive motors. The heart of 
these starters is the silicon-controlled rectifier or SCR. 
SCR's have many other applications; among these, the 
most common is in dimmers for home lighting. 

SCR's, also called thyristors, are three-terminal semi- 
conductor devices having a four-layer p-n-p-n material 
combination. Figure 5.24A shows a model of the SCR 
construction. The outer two layers act as a p-n junction 
and the inner layers serve as an element to control that 
junction. The symbol for the SCR is given in figure 5.24.B, 
and figure 5.25 illustrates how the operation of the three- 
junction combination can be equated to two transistors 
connected as shown. 

The equivalent circuit is represented by one n-p-n 
and one p-n-p transistor. When the bias on the gate, the 
n-p-n transistor base, is negative with respect to the 
cathode, the n-p-n transistor cannot conduct appreciable 
current. In other words, it is cut off. As no n-p-n transistor 
collector current can flow, the p-n-p transistor is also cut 
off. There is high impedance between the anode and 
cathode for this bias condition, and the SCR operating 
condition is called OFF. However, if the gate bias is made 
positive so that the n-p-n transistor conducts, current will 
flow into the n-p-n collector from the p-n-p transistor 



base. This p-n-p base current in turn causes collector 
current in the p-n-p transistor. The action between the 
two transistors has a positive feedback effect because an 
increase in current in one transistor creates an increase in 
the other. Therefore, once conduction in the SCR is estab- 
lished, the gate no longer has any controlling effect, and 



Metal 

Source Gate(-)/ ^ Drain 
S1O2 



AsfetfSdg 



D, drain 




p (substrate)x\\y^ 




j ga t te^^' bs ^ ate gate 



Model 



S, source 



n- channel 
symbol 



D, drain 

SS, 

substrate 

S, source 




p-channel 
symbol 



A Depletion mode operation 



D, drain D, drain 

G,o-(fs)-° SS, g, »-(jt)-^ SS, 

gate ^n substrate g a t e ^^ substrate 

S, source 



n-channel 
symbol 



S, source 

p-channel 
symbol 



B Enhancement mode operation 
Figure 5.23.— Model and symbols for MOS-FET devices. 



i> Anode 



Gate 



n — -o 



V 



7- 



Cathode 



SCR 

A B 

Figure 5.24.— SCR model (A) and symbol (8). 



-K- 



€) 



Input 



\P r 



:Ri 



R sl 4c c 4-% 



^^— 



Output 



DD 







l fl e 


c 


P 


b 


n 




n 


— o 


P 




P 




n 




C 






Figure 5.22.— Example of a junction-FET application. 



Figure 5.25.— SCR equivalent model and circuit. 



114 



the SCR is latched ON; that is, anode-to-cathode imped- 
ance becomes very low. The gate cannot turn the SCR 
conduction OFF. Cessation of current requires a negative 
gate bias and an essentially zero anode-to-cathode voltage. 
This allows the p-n-p transistor to cut off. The OFF and 
ON characteristics are apparent in the typical curve 
provided in figure 5.26. The breakover voltage noted here 
is the anode-to-cathode potential at which the SCR will 
turn itself ON. 

There are many applications for the SCR or thyristor. 
Some of the devices and system components that the 
thyristor replaces include 

• Thyratrons, 

• Mercury-arc rectifiers, 

• Saturable-core reactors, 

• Relays and contactors, 

• Rheostats and motor starters, 

• Constant-voltage transformers, 

• Autotransformers, and 

• Mechanical speed changers. 

Thyristor applications are a major subject in chapter 14. 



INTEGRATED CIRCUITS 



of this layer to interconnect the different regions. The top 
view of an actual IC is provided in figure 5.28. These 
devices can contain hundreds of transistors but can be 
small enough to pass through the eye of a needle. 

Except for high-power applications, IC's are preferred 
over discrete-component assemblies because they add reli- 
ability to equipment while reducing both size and cost. 
Consequently, IC's are employed where specific circuits 
require many transistors, diodes, and resistors. In circuit 
diagrams, it is accepted practice to show only the symbol 
for the specific application; some of these are given in 
figure 5.29. The use of IC's is extremely widespread in 
recently manufactured mining equipment, especially in 
control, monitoring, and communications applications. 



BASIC INSTRUMENTATION 

Much has been said in the preceding chapters about 
electrical parameters and their quantification: voltage, 
current, power factor, power, and so on. Instruments that 
measure these quantities are necessary to monitor and 
troubleshoot the operation of a power system and can be 
used to ensure optimum operation and to find malfunc- 
tions. The devices can be indicating instruments or record- 
ing instruments that are permanently installed in major 



The semiconductor devices discussed so far are termed 
discrete components if they are manufactured as single 
units, for example, one diode or one transistor. They must 
be combined with other electrical and electronic compo- 
nents to perform any required function. Manufacturing 
processes have been refined so that several transistors, 
diodes, and resistors can be made in a single circuit, or in 
other words on one single semiconductor chip. Such de- 
vices are termed integrated circuits (IC's), and their study 
in electrical engineering is known as microelectronics. 
Today, many circuits requiring numerous individual tran- 
sistors, such a complete amplifiers and digital computers, 
are packaged in a single semiconductor chip or microcir- 
cuit. When employing only one semiconductor chip, the IC 
is called monolithic; when the unit is created by intercon- 
necting more than one microcircuit, the device is a hybrid 
IC. 

The structure illustrated in figure 5.27 represents the 
cross section of a simple monolithic IC. The device is 
fabricated on a chip of p-type semiconductor, termed a 
substrate, by forming a number of junctions. The three 
sections shown are electrically isolated by reverse-biased 
p-n junctions, and the silicon surface is protected by a 
silicon oxide layer. A thin film of metal is deposited on top 



Avalanche 
breakdown 



ON characteristics 



U^ 



/ Breakover 
/ voltage 




OFF characteristics 



npn 
transistor 



Diffused 
resistor 




KEY 

B Base 

E Emitter 

C Collector 

n+, n 2 , n 3 Various n-type regions 



Silicon dioxide layor, metal 
film on top to interconnect 
components. 



Figure 5.27.— Sketch of simple monolithic IC cross section. 




B 30:1 enlargement of 
1 circuit 



Figure 5.26.— General characteristic curve for SCR. 



A Silicon wafer slice 

Figure 5.28.— Top view of an actual IC 



115 



30 kO 
Vjn — WW 
R 






NOR 



Amplifier within circuit 



Special-purpose circuit 
Figure 5.29.— Examples of symbols employed for IC's. 



AND NAND 

Logic circuit symbols 



equipment or they can be self-contained and portable. It is 
not unusual for every piece of power equipment in or about 
the mine to have some form of enclosed instrumentation. 
The devices can range from basic meter movements to 
transducers connected to on-line computers that monitor 
the status of the entire power-system complex. 

The word "meter" is often used as a suffix or part of a 
compound word that describes the function of the instru- 
ment. Of all the instruments designed to measure electri- 
cal quantities, the voltmeter and ammeter are the most 
basic. Voltmeters measure the potential difference or volt- 
age between two points and must present a very high 
impedance to the circuit so as not to interfere with normal 
circuit operation. Ammeters measure current flow and 
must have a near-zero impedance. The dc voltmeters and 
ammeters sense average quantities, while their ac coun- 
terparts usually provide rms voltage and current values. 
Instrument current inputs are normally at 5 A, with 
potential inputs at 120 V. 

The following section will explore the various instru- 
ments available to the mining industry, commencing with 
a description of the basic instrument or meter types and 
then showing how the devices are employed to monitor 
system quantities. 



BASIC METER MOVEMENTS 

A meter movement is an electromechanical device 
that provides the mechanical motion to an indicator in 
response to an applied electrical signal. Regardless of the 
type of meter movement, opposing magnetic fields are 
employed to activate the indicator or pointer. These move- 
ments can be classified as electrostatic, dynamometer, 
moving iron vane, and permanent-magnet moving coil. 

An electrostatic movement is the only type that 
measures voltage directly as opposed to a voltage-produced 
current. This meter is basically a variable capacitor with 
a restoring resistor connected between a fixed and a 
movable plate or vane. When a difference in potential 
exists between the plates, the opposing charges produce a 
mutual attraction and the movable vane will move toward 
the fixed vane with the deflection proportional to the 
applied voltage. Upon removal or change in potential, the 
resistor discharges the capacitance. Thus any current 
through the movement is merely incidental to the opera- 
tion. Electrostatic instruments can measure either ac or dc 
potentials; they have true rms response to ac regardless of 
waveform shape. Full-scale readings (maximum meter 



deflection) range from 100 V to 10 kV depending on the 
movement, with a measurement precision of 0.5% to 2%. 

A dynamometer movement consist of two coils, one 
fixed and the other movable. The movable coil rotates in 
the magnetic field produced by current through the sta- 
tionary coil. If the current being measured flows through 
both coils, (that is, they are in series), the resulting torque 
is proportional to the current, and the displacement is 
proportional to the square of current. Thus the pointer 
deflection indicates the rms value of current. The move- 
ment can be designed to measure dc or ac very precisely to 
within 0.1%. However, the dynamometer is not commonly 
employed as an ammeter. Its prime application is as a 
wattmeter, which, will be described shortly. 

Moving-iron-vane movements are similar to the dyna- 
mometer, except the moving coil is replaced by a soft iron 
vane with no permanent magnetization. Here, current 
through the fixed coil produces a magnetic field that induces 
magnetism in the soft iron vane. The magnetic fields oppose 
each other, producing torque that deflects the vane with a 
force proportional to the square of the current. The instru- 
ment can therefore measure dc or the rms value of ac, but 
with less precision (1% to 2%) than the dynamometer. 

The last basic type of meter movement is the 
permanent-magnet moving-coil or d'Arsonval meter, 
which is a dc ammeter. The moving element is a coil of fine 
wire suspended so that it is free to rotate in the field of a 
permanent magnet. Sketches of typical movements are 
provided in figure 5.30. When dc flows in the coil, a torque 
is produced that tends to rotate the coil. The rotation is 
opposed by some form of spring restraint, usually a helical 
spring, so that coil motion and thus pointer position is 
proportional to the coil current. If the dc through the coil 
is varying so fast that the pointer cannot follow the 
fluctuations, the pointer will assume a position relative to 
the average torque, and therefore indicate the average 
value of current. However, if the current is a sinusoid, the 
average of moving-coil torque is zero, and the pointer will 
not be deflected. Nevertheless, d'Arsonval movements can 
obtain a precision of 0.1%. 

For measuring current, both dynamometer and 
moving-iron-vane movements are often restricted to fre- 
quencies less than 200 Hz. Yet both these yield true rms 
readings within their frequency range. Electrostatic in- 
struments can be extremely precise for observing voltage, 
but they are often very delicate and are applicable only for 
laboratory use. Even though d'Arsonval movements mea- 
sure only dc, they are the most common type in use for 
both direct dc measurements and ac measurements using 
rectification. 



116 




Spring 




Pointer 
Coil 



Pivot 




External magnet 



Moving-coil construction 

Figure 5.30.— Permanent-magnet moving-coil movements. 



Internal magnet 



Meter-Movement Applications 

When a d'Arsonval meter is used as an ammeter, it is 
inserted in series with the circuit being measured. The 
current range for this direct application is obviously 
restricted by the maximum scale reading or maximum 
current of the movement. D'Arsonval meters can have 
full-scale limits from 1.0 /zA to 50 mA, although the basic 
movement is considered to be 1.0 mA, which allows 
measurement from zero to 1.0 mA. For higher current 
requirements, the meter is shunted with a low resistance 
as shown in figure 5.31. Such shunts can be tapped to 
provide several current ranges, or several shunts might be 
available, each selected by a switch to provide a specific 
current range. Commercially available ammeters of this 
type offer up to a 50-A full-scale reading. 

To measure dc voltages, a d'Arsonval movement is 
simply placed in series with a selected high resistance, 
and the combination is connected between the two points 
where a voltage measurement is desired (fig. 5.32). Be- 
cause meter deflection is still proportional to current, the 
meter scale can be calibrated to read the voltage required 
to produce a specific current. The sensitivity of such 
voltmeters is stated in ohms per volt. For instance, if a 
meter has a range of to 200 /zA and if the movement is to 
be used to measure to 200 V, the total meter resistance 
must be 



R = 



200 V 
200 M A 



= 1.0 Mfi. 



As moving-coil resistance, R m , is generally on the order of 
50 to 100 12, it can be neglected in this case. Sensitivity of 
the combination is therefore 



1,000,000 fl 
200 V 



= 5,000 Q/V. 



A higher value of sensitivity for a specific meter implies 
higher quality. Presently, the upper limit for the commer- 
cially available d'Arsonval voltmeter is 50 k(2/V The 
standard d'Arsonval movement of to 1 mA has a coil 
resistance of 100 Q; hence, it can be employed to read to 
100 mV directly. 

External shunts are utilized for a desired maximum 
current when the current is higher than measurable by 



Permanent 
magnet 

RM = 100 n, 0-100 mV, 0-1 mA 
Load line /^\ 

1mA 

Shunt 




Figure 5.31.— Shunting d'Arsonval meter for high-current 
tests. 



• Permanent 
magnet 



RM=100il 
0-200 uA 




Figure 5.32.— D'Arsonval meter used to measure dc poten- 
tials. 



normal instruments with internal shunts. Figure 5.33 
provides a couple of typical constructions where terminals 
are available for circuit as well as meter connections. 
These are simply standard resistance units, designed to be 
used with either 50-mV (0- to 50-mA) or 100-mV (0- to 
100-mA) movements, in which a current through the 
shunts is indicated by a specific voltage drop across the 
shunt. For example, if a shunt is designated 100 mV, 600 
A, a reading of 50 mV across the shunt signifies that 300 
A is flowing in the circuit. Any time that metering or 
instrumentation is part of dc mine power equipment, it 
can almost be assumed that external shunts are involved. 



117 



lb this point, only the measurement of circuit opera- 
tion has been considered. A d'Arsonval meter can also be 
used to measure resistance by the addition of a dc source in 
the dc voltmeter circuit. Consider the circuit shown in 
figure 5.34, which has a dc movement in series with a dc 
source (usually a battery) and one or more resistors, one of 
which is usually variable to be used for calibration. The 
unknown resistance to be measured completes the loop. 
Meter deflection is still proportional to dc through the loop 
and is therefore a function of the unknown resistance. 
Using known resistances, the meter scale can be cali- 
brated to read resistance directly, and different fixed 
resistors or multipliers can be used to extend the single 
scale. The combination is easily calibrated before each use 
by adjusting the pointer to zero using the variable resis- 
tance. The resistance desired could be a simple component 
or a complex circuit, but the ohmmeter should never be 
used on an energized circuit because of the internal 
source. 

Combining the d'Arsonval movement with a half- 
wave or full-wave rectifier allows the reading of ac values 
in terms of dc through the coil. The full-wave or rectifier- 
ammeter circuit shown in figure 5.35 is the most common. 
Here, current through the movement is I d , and thus, meter 
deflection is proportional to the average of I d . This reading 
is the half-cycle average if the ac is symmetrical (that is, 
the dc scale of the meter will read the half-cycle average 
sinusoidal current). As the rms value of current is usually 
desired, the scale is calibrated in rms by multiplying the 
average current by 1.11. This is the rms value for a 
sinusoidal waveform only; for any other waveshape, rely- 
ing on the rectifier circuit can produce large errors. 

Moving-iron and dynamometer movements record rms 
current automatically, and many permanent meters built 
into power equipment to measure ac voltage and current 
are moving-iron types. However, the d'Arsonval meters are 
often preferred because of their greater sensitivity. For ac 
measurements of voltage or high current, the concepts of 
high series resistance and low parallel resistance also can 
be applied to the rectifier, moving-iron, and dynamometer 
movements, but such practices are not common except in 
small portable test equipment. 

It can be seen in the foregoing that the d'Arsonval 
meter is used to measure ac or dc voltage or current as well 
as resistance. An instrument incorporating all these func- 
tions is called a multimeter. The selection of a specific 
parallel or series resistance combination provides the 
needed measurement function and parameter range. 



Line terminal 



Instrument 
terminals 




Line terminals 

^Instrument terminals 
Copper blocks 

Line terminal 



Figure 5.33.— External shunts used for high-current 
measurements. 




Unknown 
resistance 



Figure 5.34.— Simple ohmmeter circuit. 





i„A A ( 
Q A A 
r. VWYY 



Figure 5.35.— Rectifier ammeter. 



Wattmeters 

As mentioned earlier, the main application for dyna- 
mometer movements is in wattmeters. Figure 5.36 illus- 
trates the wattmeter connection. Typically, the fixed coil 
carries circuit current while the moving coil is connected 
in series with a high resistance and is attached across the 
terminals of the circuit (the moving coil can itself be of 
high resistance). Circuit current flows through the fixed 
(or current) coil, and the current through the moving (or 
potential) coil is proportional to circuit voltage. Therefore, 
the movement torque is proportional to the product of 
instantaneous voltage and current, with the indication 
relative to the produce average or average power. The 
dynamometer connected as such will measure correctly 
the average power of a dc or ac circuit of any waveform, 
even when a power factor is involved. 




Current 
coi 



Current 
coil 



r o Loads 6 



Figure 5.36.— Dynamometer connected as wattmeter. 



118 



Varmeters 

In addition to being used for measuring watts, the 
dynamometer movement has wide application in measur- 
ing reactive power or vars. This is done in single-phase 
instruments by shifting the phase of the voltage coil by 
90°. The voltage coil flux is then in phase with the flux 
produced by the reactive-current component in the current 
coil. Varmeters are installed in the same manner as 
wattmeters are. 

Power-Factor Meters 

A power-factor meter shows the power factor continu- 
ously and indicates whether the current is leading or 
lagging the voltage. The movement resembles a single- 
phase wattmeter but has no control spring and has two 
moving potential coils mounted on the same shaft 90° 
apart. One potential coil (B of figure 5.37) is in series with 
a noninductive resistor so that it produces torque propor- 
tional to the line voltage and in phase with the real 
component of line current. The other coil (coil A) is in 
series with a higher quality inductance, so its torque is 
proportional to the line-current reactive component. The 
fixed coil (coil C) is again the current coil. With unity 
power factor, the average torque between coils A and C is 
zero since the currents are 90° apart, but the currents 
through coils B_an_d C are in phase, so the torque p roduced 
aligns their axes, and the pointer indicates unity power 
factor (1.0 pf). For leading or lagging power factors, the net 
torque created by currents in coils A, B, and C will swing 
the moving coils to the right or left, aligning the pointer in 
a position relative to the power factor. Meter scales are 
therefore calibrated so that the center position is unity 
power factor, and to the left and right of center are lagging 
and leading power factors from unity to zero. 

This section has presented some direct applications for 
basic meter movements. Some concepts shown here apply to 
all electrical parameter measurements, but for ac power 
systems, additional components are normally employed. 

POWER-SYSTEM INSTRUMENTATION 

In chapter 3, the subject of current transformers (CT's) 
and potential transformers (PT's) was introduced. These 
devices actually fall under the general category of instru- 
ment transformers and serve two main functions: 

• To isolate instruments, relays, and meters from line 
voltage, and 

• To transform line currents and voltages into values 
suitable for measurement by standard instruments. 

Thus, the normal ratings of instrument transformer sec- 
ondaries are 5.0 A for CT's and 120 V for PT's. This 
measurement implies not only metering or actual visual 
readings but also sensing for such purposes as protective 
relaying. The following material will cover specifics of 
CT's and PT's as they apply to instrumentation of mine 
power systems. Chapter 9 will discuss the application of 
these transformers to protective relaying. 

Instrument Transformers 

Instrument transformers are connected in the power 
system in a manner related to the function they monitor. 
The primary winding of a CT is placed in series with the 




Load 



Figure 5.37.— Power-factor movement. 



9- 




Source 



H 1 H 2 Load 



<!— 


CT 


\ 


HHHf- 

*1 


j -* 

■x 2 








Hi" 











Ammeter 

Voltage Current 

Figure 5.38.— Simple instrument-transformer connections. 



line conductor to be measured, or may be the line conduc- 
tor itself, while a PT is placed across the line voltage to be 
measured (fig. 5.38). The transformers can then be used to 
extend the application of ac instruments in the same way 
that shunts and series resistors extend dc instrument 
usage. In this case, the ratio of a CT or PT is the ratio of 
primary current or voltage to secondary current or voltage 
under specified conditions. The secondary winding param- 
eter is coordinated with the connected instrumentation. 

To operate reliably, an instrument must receive infor- 
mation that accurately represents the conditions existing 
on the power system. When operated outside of the range 
for which they are intended, instrument transformers are 
very nonlinear devices; that is, the output from the trans- 
former secondary can deviate from being an accurate 
representation of primary-winding conditions. The 
amount of deviation is a function of the transformer input 
level, secondary load, and design. To help with current 
application of instrument transformers so that they oper- 
ate in their linear range, the American National Stan- 
dards Institute (ANSI) has standardized transformer de- 
signs and secondary loads. 1 The designs are called 
accuracy classes, and the secondary load is called the 
transformer burden. 

The effects of burden changes are typically more 
pronounced with CT's and PT's. Preferably, CT burden is 
expressed as a standard load impedance or its resistance 
and reactance components. In the past the practice was to 
specify the value as an apparent power (in voltamperes) at 
a power factor, the angle of which was referenced to a rated 
secondary current (for example, 0.9 pf of current lagging). 
Consequently, a CT burden of 0.5-fl impedance could be 



1 Requirements for Instrument Transformers. C57.13 1968 et. seq. 



119 



expressed as 12.5 VA at 5 A, assuming the usual 5-A 
current. However, because of the nonlinear nature of 
transformers, burden impedance decreases as the second- 
ary current increases, and a specific burden may apply 
only to one level of secondary current. As a result, the now 
nonstandard voltampere ratings are confusing. Further- 
more, CT burden must be applied not only to the external 
load but also to all elements of that load, including the 
interconnecting loads. As the total burden needs to be 
calculated frequently, manufacturer publications usually 
provide the burdens of individual components. Potential 
transformer burden is normally stated as the total exter- 
nal voltampere load on the secondary at rated secondary 
voltage. 

For the best accuracy with either PT's or CT's, the 
impedance of the burden should be identical to that of the 
instrumentation, and the accuracy limits stated by ANSI 
will then apply. The general rule for CT's is that if silicon 
steel is used for the core, the ampere turns should be at 
least 1,000 for good accuracy under normal conditions. 
When a PT has acceptable accuracy at its rated voltage, it 
can normally be used over a range from zero to 110% of 
rated voltage. Operation greater than 10% overvoltage can 
produce excessive errors. 

Some special precautions are in order whenever cur- 
rent transformers are in use. A CT secondary should 
always be shorted or properly connected to the instrumen- 
tation (meters, relays, etc.), or dangerous potentials can 
occur at the secondary terminals and the core can become 
permanently magnetized. The flux density in the core is 
normally very low and can rise to saturation without a 
secondary current. The core can also become magnetized if 
dc is passed through the secondary. In either situation, the 
transformer ratio can be seriously changed. Furthermore, 
it is possible for a CT to be damaged through insulation 
breakdown associated with surges, overloads, and other 
occurrences. Therefore good practice dictates that tests be 
conducted prior to installation and periodically thereafter 
to verify transformer operation. If magnetization is sus- 
pected, the core can be demagnetized by passing rated 
60-Hz current through the secondary with the primary 
open and gradually reducing the current to zero. 

When a fault occurs on a line downstream from the 
CT coupling, the transformer primary current may reach 
several times the rated value for short periods of time. Two 
different techniques are available to protect against CT 
damage. One method is to overdesign the primary winding 
so that the transformer will not be damaged by the 
mechanical and thermal effects of moderate overload. The 
other design is perhaps more desirable. Here the CT is 
selected so that its core is close to the saturation point 
with normal operating primary current. When a surge 
current occurs, the secondary current cannot increase in 
proportion to the primary current and the burden is thus 
spared much of the shock. (See chapters 9 and 10 for 
further details as core saturation can seriously affect 
protective-relay operation.) 

Single-Phase Connections 

Figure 5.39 illustrates the measuring device connec- 
tions needed for a single-phase circuit in order to observe 
voltage, current, and average power. This is a simple 
extension of figure 5.38. Only two instrument transform- 
ers are required: the PT drives the voltmeter and the 
wattmeter voltage coil, and the CT supplies current to the 



ammeter and the wattmeter current coil. For this arrange- 
ment, the ammeter and voltmeter would probably be 
moving-iron movements and the wattmeter would be a 
dynamometer. An alternative instrument arrangement is 
illustrated in figure 5.40. Here transducers are placed 
between the instrument transformers and the meter move- 
ments. Transducers are electronic components that 
present a standard burden to the transformer and provide 
an output compatible with the standard d'Arsonval move- 
ment. This is usually to 1 mA, but to 50 mV and to 
100 mV are also available. The transducer output is also 
adapted to a range of load impedances. With either ar- 
rangement, three ac power parameters can be measured 
and the power factor can also be calculated if desired. 
When any meter movement is employed, the normal 
reading of the meter should be one-half to three-quarters 
of the full-scale value in order to provide the best precision. 
Note that in figures 5.39 and 5.40 the instrument 
transformer secondaries are grounded. The grounding is 
needed to prevent a high static potential, which can cause 



Power 
conductors 



To loads 




Voltmeter 



Ammeter 



Figure 5.39.— Voltmeter, ammeter, and wattmeter arranged 
as single-phase system. 



nonn 
5.0A 



To loads 



PT 



120 V 



Voltage 
input 



0-120V 

Voltage 

transducer 

0-1 mA 



Output 



Power 
transducer 

<U 
o 

i 
o 

mA 



> 

o 

£^ 

i 
o 

O- 



Output 




Current 
input 



% 



CT 



0-5.0 A 

Current 

transducer 

O-l mA 



Output 



0-1 m A meters 



Figure 5.40.— Use of transducers with standard d'Arsonval 
movements. 



120 



a higher voltage than normal to appear on the secondar- 
ies. Without grounding, the transformer insulation could 
fail. The transformer case should also be grounded for the 
same safety reason. 

Three-Phase Connections 

When the measurement of average power in a three- 
phase system is required, it seems obvious to place one 
dynamometer wattmeter in each phase and add the re- 
sults together. This is shown in figures 5.41A and 5.41S 
for a four-wire wye load and a three- wire wye or delta load. 
The sum of the meter readings is total power for either 
connection, for any waveform, and whether the system is 
balanced or not. The common connection of the three 
wattmeter potential coils may be placed at any potential 
without affecting the total power readings. If the potential 
is that of one phase conductor (see figure 5.42), one 
wattmeter becomes inoperative and thus may be omitted. 
The result is the two-wattmeter method of three-phase 
power measurements. Commercially available transduc- 
ers can be used instead of the two wattmeters. The 
transducer inputs are two line-to-line voltages and two 
line currents, and the single output, which is proportional 
to total power as before, can be used with a standard 
d'Arsonval movement. A circuit arrangement for this 
method is shown in figure 5.43. 

Under balanced conditions, the readings from the 
two-wattmeter method can be used not only for total power 
but also to determine the power-factor angle. It can be 
shown that 



Current Wattmeters 

connection _ / 



Voltage 
coil — j=r 



tM 





Figure 5.41.— Three-phase wattmeter connections. 



Wattmeter Clirrpnt rnil 



Potential 
coil 



Source » 




• Load 



Potential 
coil 



Wattmeter Current coil 
Figure 5.42.— Two-wattmeter method. 



tan = V3 



P2-P1 
P 2 + Pi ' 



(5.14) 



where P 1; P 2 = two power readings, corresponding to ar- 
rangement in figure 5.42, 
and 6 = load power-factor angle. 

If P x represents a measurement of phase a current, equa- 
tion 5.14 provides the correct sign for the power-factor 
angle, thereby specifying whether the load is capacitive or 
inductive. At times, phase sequence is hard to distinguish 
in practice, but the equation yields the angle magnitude 
and this is often sufficient information since the reactive 
characteristics of the load are usually known. 

If the system is balanced or can be approximated as 
such, the circuit shown in figure 5.44 can be employed to 
measure the line-to-line voltage, line current, power factor, 
and total average power. The two-wattmeter approach 
calls for two PT's and two CT's. One PT supplies the 
voltmeter and one CT provides information to the amme- 
ter, while the remaining PT and CT supply the power- 
factor meter so that the transformer burdens are balanced. 

It is often useful to observe each line current or 
line-to-line voltage for major power equipment. Figure 
5.45A provides an economical method for the line currents 
in which only two CT's are needed. If one CT secondary is 
measured, the current will correspond to the CT phase 
(that is, phase a or phase c), but if both CT secondaries are 
in parallel, the current reading is for the phase without 
the CT (that is, phase b). This metering is theoretically 
correct only for balanced voltages, but on most systems the 
voltage is close enough to balance that the two-CT ap- 
proach gives acceptable precision. If greater accuracy is 



Source 



Load 




To standard 
0-1 mA meter 



Figure 5.43.— Three-phase power measurement with 
transducer. 



needed, three CT's should be used as shown in figure 
5.45S. It is possible to connect the CT secondaries in delta 
or wye, but the burden impedances should always be wye 
connected. To observe all three line-to-line voltages, three 
potential transformers can be used as in figure 5.46A. The 
open-delta arrangement shown in figure 5.46S is not as 
accurate but gives satisfactory precision and uses only two 
PT's. For current or voltage with two or three instrument 
transformers, power-equipment metering is performed 
with a voltmeter or ammeter or both. The required phase 
is switch selected by connecting the transformer combina- 
tion to the meter. 



Source 



Yyyyyy^ 



^npnpn 



Neutral 



6 High 



Med 



Low 



• 6 



• 6 



oHigh 



• +4+3+2 + 
Low 



Med 
-o 



Low 



Med 



High 



• +4 + 3+2+1 
Low< 



Med 




t 



ti=± 



u 



Watts 



+ t t 

Power-factor 
meter output 



High 



Voltage 



i C C "»7 In 1 

u '»8 2"—j J 

I -10 4*'- I t t < C * 

r— »I1 5>" C C ' 

"12 6"-£-£ ' 



Load 



4 3 2 1 



Fl 



Current 



4 3 2 1 



m 



nxt 



Average - power 
meter output 



Voltage Current 
meter meter 
output output 



Figure 5.44.— Balanced three-phase measurement of voltage, current, and average power. 



121 



Source 




Meters n^\ n 



•-Load 




*- Load 



Source 




Load 




■*- Load 



Meters 



B 



Figure 5.45.— Line current measurements with two or three 
CTs. 



Figure 5.46.— Line-to-line voltage measurements with three 
or two PT's. 



122 



SPECIAL INSTRUMENTS 

Several special, if not very common, instruments are 
available to perform measurements on specific electrical 
quantities. These include but are not limited to watthour 
meters, demand meters, bridges, megohmmeters, and 
phase-sequence indicators. Each of these is described in 
the following paragraphs. 

Watthour Meters 

The watthour meter is a common power instrument, 
used in nearly every building to measure consumed elec- 
trical energy. The typical watthour meter consists of a 
small induction motor with an aluminum disk that is 
rotated by a torque proportional to voltage times current 
at every instant. The principle of operation is similar to 
that of the dynamometer wattmeter, except the disk is 
allowed to turn continually with a speed proportional to 
average power. The number of turns is counted by a train 
of clocklike gears. The counter thus indicates the product 
of power and time, or energy, which is measured in 
kilowatthours. A simplified sketch of the induction mech- 
anism is shown in figure 5.47. 

Demand Meters 

Demand meters are usually of two types (although 
there are others): integrated demand or lagged demand. 
The readings may be indicating or recording. Integrated- 
demand meters consist of an integrating meter element, 
such as the watthour meter just described, that totals the 
energy used over the demand interval and drives a maxi- 
mum indicating device, which can be a passive pointer, 
display, or chart. The meter can be reset manually, or a 
timing device can be used to return the drive to zero at the 
end of the recording period, thus leaving an indication of 
maximum demand. Lagged-demand meters provide a 
maximum demand indication that can be subjected to a 
characteristic time lag by either mechanical or thermal 
means, but usually the exponential heating curve of 
electrical equipment is followed. The demand interval is 
then defined as the time required to indicate 90% of the 
maximum value of a suddenly applied steady load; thus, 
maximum demand can be observed. Demand meters, 
whatever the type, can provide input to the power-system 
studies. 

Bridges 

Bridge circuits yield the most precise measurements 
of impedance, be it resistance, capacitance, or inductance, 
for two reasons: the measurements rely on null methods, 
and comparisons are made directly with standardized 
impedances that are precisely known. The term null 
method means that a zero reading or null indicates the 
correct value. 

The Wheatstone bridge is the most widely used of 
these circuits. Shown in figure 5.48, the bridge is dedi- 
cated to measuring resistance, capacitance, or inductance 
depending on its internal components. 

When the Wheatstone bridge is intended to measure 
resistance (figure 5.48A), the circuit consists of two fixed 
precision resistances, R 1 and R 3 , which are known as the 
ratio arm; a variable precision resistance, R 2 ; and the 



, Transformers 



Dial 




Line 



Permanent magnets 



Aluminum disk 



Simple schematic 




Line 



Load 



Eddy currents produced Voltage coil 

by voltage coil (highly reactive) 

Disk plan view 

Figure 5.47.— Simplified sketch of watthour meter induction 
mechanism. 



Galvanometer 




Unknown 




A Wheatstone bridge for resistance 



Audible device 
or meter 




Unknown 



B Impedance measurements with a Wheatstone bridge 
Figure 5.48.— Wheatstone bridge circuits. 



Unknown 



123 



unknown, R x . A dc source supplies current to the arrange- 
ment, and a galvanometer, G, is located at the center of the 
bridge across points b and d. The galvanometer is simply 
a very sensitive ammeter with a center-scale zero-reading 
pointer and the ability to read very small currents in 
either direction. R 2 is adjusted to provide a null reading on 
the galvanometer, which means the potential between b 
and d must be zero. With this balanced condition, the 
unknown resistance can be calculated by 



R x = ^~ Ra 



(5.15) 



In commercially available bridges, R x , R 2 , and R 3 are all 
variable and the value of each is readily determined by 
calibrated dials. Thus, the bridge can measure resistances 
precisely over a broad range. 

To measure impedance, R 3 of the resistance bridge is 
replaced by Z 3 , and the unknown is now Z x . An ac source 
is used, together with some means of measuring the 
potential between points b and d. This could be a sensitive 
ac ammeter or an audible device such as a set of head- 
phones. R x and R 2 are then adjusted to provide a null, and 
the balanced condition means that 



z, = z s | 



Megohmmeters 

The preceding resistance-measuring devices can be 
ineffective when resistance is in the many millions of 
ohms. An important factor here is the resistance of insu- 
lation, such as that around conductors (fig. 5.50). One 
problem in these and other high-resistance measurements 
is to provide sufficient potential so the resulting current 
can be detected by an indicating device that provides 
resistance readings. The instrument designed to perform 
these tests is called a megohmmeter (fig. 5.51), where the 
unknown resistance is R x , and R x and R 2 serve as current- 
limiting resistors to protect the meter from damage. 



Galvanometer 



(5.16) 




Obviously, the values of Z x and Z 3 depend upon the 
frequency of the ac source. The most typical value used is 
1,000 Hz. 

If very low resistances in the order of 10 jtQ to 1.0 mfl 
must be measured, the Kelvin double bridge shown in 
figure 5.49 can be used. The circuit consists of ratio arms 
R A , R B and R a , R^; a connecting link or conductor, R f ; a 
known resistance, R s ; the unknown, R x ; an adjustable dc 
source; and a null indicator. The indicator could again be 
a galvanometer. The resistances r x , r 2 , r 3 , and r 4 are those 
of the connecting leads between the four-terminal bridge 
and the resistances to be compared (R x and R 8 ). These lead 
resistances should be in the same ratio as the bridge arms 
to which they are connected; otherwise, the ratio unbal- 
ance will cause incorrect measurements. A small adjust- 
able resistor can be used to balance the lead resistances. 
The balance equation is thus 



Figure 5.49.— Kelvin double bridge. 



R_x 
R„ 



R, 
R T 



R s \R a + Rg + R f / VRg R«/ 



Instrument 
test leads 



Conductor 
insulation 




Indicating scale 
shows resistance 



Megohmmeter 
Figure 5.50.— Megohmmeter testing insulation resistance. 



When R x and R s are so small that R f is comparable, the 



term in equation 5.17 involving R f 
However, if 



can be significant. 



R, 



Rfl 



then the R e term becomes zero. The source is adjustable so 
that current through R x , R e , and R 8 (the series resistance 
of which is small in comparison to the bridge) is large 
enough to allow a measurable milling current through the 
indicating device, G. An application for the Kelvin double 
bridge is in the measurement of cable and conductor 
resistances. 




Generator- C_Hand crank 
Figure 5.51.— Internal components of megohmmeter. 



124 



The most evident difference between the megohmme- 
ter and the preceding instruments is the hand-driven 
generator, which supplies the needed dc potential for 
measurement. The generator applies from 500 to 2,500 V 
depending on the instrument and is tied to the resistance 
range desired (the higher the measured resistance, the 
higher the required voltage). Typically, a friction clutch is 
employed to restrict the generator to rated output voltage. 
In some megohmmeters, the potential is from batteries via 
an electronic power supply located within the instrument. 

As shown in figure 5.51, the meter has two coils 
mounted over a gapped core. The movement is similar to 
the d'Arsonval, but there are no restraining springs, so the 
indicator is free to move when there is no output from the 
generator. If the instrument terminals are open (that is, R x 
is infinite) when the generator is operated, current will 
flow through R 2 and coil A x , and the torque produced will 
force the pointer counterclockwise to the infinite scale 
reading. When the terminals are shorted (R x is zero), the 
torque produced by coil B is greater than that from coil A 
and this moves the pointer to a zero reading. For measur- 
ing an unknown resistance, the pointer location is depen- 
dent upon the opposing torque from the two coils, and the 
position is a function of R x . 

Another prime application for megohmmeters is the 
measurement of ground-bed resistances. These specialized 
testing procedures are covered in chapter 7. 

Phase-Sequence Indicators 

In order to prevent damage or incorrect operation, all 
conductors in a three-phase distribution system must be 
properly connected so they will provide the same phase 
sequence to all equipment. Correct interconnections can at 
times be difficult to accomplish in mine power systems, 
especially with feeder and trailing cables. At present there 
is no standard color coding for phase conductors. The 
phase-sequence indicator illustrated in figure 5.52 can be 
used to determine the phase relationship of energized 
three-phase conductors. It falls in the simplest class of 
testing devices: indicating instruments; other examples 
include a light bulb with leads to test for the presence of 
potential, or a battery in series with a light bulb with 
leads to check continuity by completing the series circuit. 
The phase-sequence indicator consists of two light bulbs 
and a capacitor connected in wye, and the lamps are 
labeled in the two possible phase combinations. Because of 
this arrangement, one lamp will burn brighter than the 
other depending on the connections to the power system. 

The foregoing has provided information on several 
devices that are helpful in measuring mine electrical 
systems. Other instruments that are equally useful for 
specific applications include the split-core ac ammeter, a 
handheld ac ammeter that has its own CT; the synchro- 
scope, which measures proper phase connections and the 
correct speed of parallel ac generators; and a frequency 
meter, which indicates the frequency of an electrical 
supply in hertz. Often there is also a need to obtain a 
continuous record of an electrical parameter, and the next 
section discusses the popular recording devices. 



RECORDING INSTRUMENTS 

Many of the direct-reading indicating instruments 
just presented are also available as recording instruments. 



Some of these are very similar to their indicating counter- 
parts in that they can use the same electrical movements; 
they differ because the pointer is also used to provide a 
graphic record on a chart. These are termed chart record- 
ers; one popular class is strip-chart recorders, so named 
because the electrical parameter is recorded on a strip of 
paper. 

The similarity between the movement of the strip- 
chart recorder and the indicating instruments is illus- 
trated in figure 5.53. The strip-chart recorder movement is 
actually a d'Arsonval type. The pen can trace on paper in 
several ways. 

• Inking. The pen is a capillary tube through which 
ink flows from a well to the chart. This is perhaps the most 
used system. 

• Inkless. The tip of the pen is a stylus that impacts 
the paper like a typewriter key with a regular force 
supplied by a cam, leaving a series of dots. 

• Thermal. The pen tip contains a heating element 
that leaves a trace by heating specially treated paper. 



Feeder line 




Figure 5.52.— Phase-sequence indicator. 



Input circuitry to 

condition input voltage 
and establish sensitivity 
of the recorder 



Comparator, compares input with 
reference, outputs a voltage in 
proportion to the needed 
position of the servomotor 



Amplifier, amplifies 



Circuitry to 
condition pen- 
sensor output 
and establish 
reference, 
including to 
zero setting 



Paper strip 
chart, driven 
at varbus 
constant speeds 




y comparator output to 
drive servomotor 



Servomotor, 
drives mechanical 
pen system 



Indicator and inking pen 



Figure 5.53.— Strip-chart recorder. 



125 



The simplest unit provides a curved recording as the pen 
swings in an arc, but articulated pen arms are also 
available that produce linear or rectilinear traces. The 
paper chart moves past the pen at a predetermined speed 
driven by an electric motor or a mechanical-spring clock- 
work mechanism. This recorder provides a continuous 
record of the average or rms value of the electrical param- 
eter of interest, which is advantageous in obtaining 
records of equipment operation, for example, the electrical 
performance of a mining machine. A variation of these 
recorders uses a round chart, driven like a disk on a record 
player but at very slow speed. These charts can be built 
into major equipment to provide permanent records. 

Sometimes recordings of the actual electrical wave- 
forms are needed to study power systems. This calls for an 
instrument that can resolve instantaneous values of elec- 
trical parameters. Electromechanical instruments that 
have this resolution are called oscillographs, and the 
movement in most of these is a sensitive galvanometer of 
low mass. Two types of writing systems are normally 
available: 

• Direct writing. This is similar to either the inking 
or thermal strip-chart recorder types. The pen has high 
inertia, and instrument response is about 0.5 to 100 Hz 
(some to dc). 

• Optical. Instead of a pen, the movement drives a 
low-mass mirror that deflects a light beam that exposes a 
light-sensitive paper. Developing is required to obtain the 
record, but the system can have resolution to 10,000 Hz. 

For many applications, magnetic tape recorders and oscil- 
loscopes, both electronic instruments, find favor over oscil- 
lographs. However, oscillographs still have some practical 
use, especially where an extended-time hard copy is 
needed immediately. An example would be in measuring 
neutral currents existing on three-phase equipment, 
which can have dc as well as ac components. 



ELECTRONIC INSTRUMENTS 

The employment of complex and sophisticated control 
equipment in the mining industry is continuing to in- 
crease. Instances include solid-state motor starters, elec- 
tronic protective relaying, computer logic circuits on min- 
ing machinery, and so on. These types of systems require 
precise voltage, current, and waveform measurements 
that are not possible with the preceding instruments. 
Certain phenomena existing on power systems, such as 
transients, require precise measurements with frequency 
response into the megahertz. Electronic measuring equip- 
ment answers this need. This section will introduce only 
the more popular instruments. 

Electronic Meters 

These instruments use many of the basic circuits that 
have been described for multimeters; that is, series resis- 
tances for voltage (fig. 5.54A), voltage-drop for resistance 
(fig. 5.545), and shunts for current. The prime difference is 
that a scaled-down dc voltage, which is proportional to the 
actual circuit voltage, current, or resistance, is amplified 
by electronics. When the parameter is sinusoid, the ac is 
rectified before amplification. The amplified signal then 
drives the indicating device. In the past, vacuum tubes 



Range switch 
(attenuator), 
allows selection 
of sensitivity 



Filter, removes 

any ac superimposed 

on dc 

Amplifier, amplifies 
dc input drive meter 



Multiplier 
resistors 




Meter 



Feedback, stabilizes 
amplifier characteristics 



Full-scale sensitivity of voltmeter 
A dc voltmeter 



Range, value per division 
on meter scale -\ 



Resistance 

being 

measured 



^1,000 n 



Multiplier 
resistors 



IOO n 



10fl<? 



Amplifier, amplifies dc voltage 
drop across unknown resistance, 
which is proportional to its 
ohmic value 



in 



\ 

Range T 
switch' — 
Battery, acting 
t as current source 
:±i through measured 
resistance 



-^ — I 



Meter, 



Calibration CV) calibrated 



resistor 



i 



in ohms 



B Ohmmeter 
Figure 5.54.— Input circuits on electronic voltmeter. 



performed the amplification, termed a vacuum-tube volt- 
meter or VTVM; more recently, solid-state devices (IC's or 
FET's) have become the most popular. 

The indicating device can be of two types: the familiar 
d'Arsonval movement or a digital indicator. The electro- 
mechanical displays or movements described thus far can 
be termed analog. The digital display is an indicating 
output assembly that takes the measurement results 
(voltage, current, average power, etc.) and through elec- 
tronics gives a visual indication in a discrete number, as 
shown in figure 5.55. The actual display can be by Nixie 
tube (a gas-discharge tube), seven-segment incandescent 
filament, light-emitting diodes (LED), or liquid-crystal 
display (LCD). The electronics in the display assembly use 
logic or binary mathematics to convert the analog output 
of the measurements and drive the visual display. These 
digital displays are replacing their analog counterparts in 
many applications. 

Electronic meters might appear rather complicated, 
but an important advantage is gained through the cir- 
cuitry: the instrument can be made so that its interference 
with the circuit being measured is negligible. Typical 
input impedance of most electronic voltmeters is 11 MQ. 

Oscilloscopes 

Oscilloscopes are electronic instruments that provide 
a real-time display of waveforms. They are available with 
responses from dc to hundreds of megahertz and thus can 



126 



ANALOG-TO -DIGITAL CONVERTER, 
generates time interval in proportion 
to input voltage 



o- 



Input 
attenuation 



Comparators compare 
input voltage with 
reference; when the 
voltage at + terminal 
is greater than at 
the — terminal, the 
output of that device 
changes from zero 
to a positive voltage; 
the 2 important 
comparison points 
are zero (for start 
pulse) and the input 
voltage magnitude 
(stop pulse) 



Input 
comparator 

/ 

-♦ » 



Generates 
stop pulse 
with 2 input 
voltages . 



J5 



ANDjf- 



\ 

Zero- 

voltage 
comparator 



AND V- 



Generates 
start pulse 
with 2 input 
voltages I 
l 



Counter, measures time interval and 
outputs a sequence of pulses to display 
the result, at a high rate such that 
display appears continually 



Amplifiers, amplify 
counter output for 
display 



Reset 



Time - 

interval 

stop 



Time- 
interval 
start 

Reset 



Reference (ramp) 
generator, produces 
an increasing voltage 
waveform to be 
compared with the 
input 



Sampling-rate 
generator, establishes 
the starting point for 
each measurement and 
the time between 
consecutive 
measurements 



Display- 
segment 
amplifier 



-*\ Digits 
amplifier 



1.1.1 



mil 



1:1:1 



Display (7-segment devices are shown); 
output pulses from counter activate 
proper digit and segments of that 
digit, then another digit and its proper 
segments, sequentially until the 
number is displayed 



Figure 5.55.— Digital display. 



be used to observe a large range of electrical phenomena 
including those of extremely short duration. The reason 
these instruments have such a broad frequency range is 
that they are not constrained by mechanical inertia. The 
heart of the oscilloscope is a cathode-ray tube or CRT (fig. 
5.56). A fine beam of electrons is deflected by an electro- 
static field in relationship to the voltage or current being 
investigated. The beam then impinges on a fluorescent 
screen to create a luminous display. The electrostatic field 
is normally established by two pairs of deflecting plates; 
one provides deflection vertically, the other horizontally. 
When a waveform is observed, the horizontal pair is driven 
electronically by a sweep signal to provide a time base, and 
the vertical pair creates a field in response to the instan- 
taneous value of the electrical parameter. In some CRT's, 
additional pairs of vertical plates are available that enable 
more than one trace to be displayed on the screen. This 
allows direct comparison of two or more waveforms. A 
camera can be used in conjunction with the oscilloscope to 
provide a permanent record. 

Tape Recorders 

The familiar magnetic tape recorder records a signal 
by magnetizing a thin strip of tape. The nonmetallic tape 
is coated with a very thin layer of magnetic material such 
as iron oxide, thereby providing a relatively permanent 
record of a signal. The signal can be an analog recording, 
or digital, or in direct relationship to the measured param- 
eter. With the digital recording, the signal is converted 



electronically to the binary system and the binary coun- 
terpart is recorded. The recording can be played back 
numerous times, and the output very closely matches the 
input that was observed. Analog tapes, which can have 
frequency ranges from dc to over 20,000 Hz, can be used as 
input to strip-chart recorders to provide hard copies or 
input to various electronic instruments that perform anal- 
ysis of the electrical parameters. Digital recordings can be 
made compatible with digital computers for swift and 
elaborate analysis of the data. 

Transducers 

Transducers perform the transfer of information from 
the power system to the electronic instruments. A trans- 
ducer can be described as a device that provides an 
electrical signal output in response to a specific measure- 
ment. Thus, potential transformers and series-dropping 
resistors can be considered voltage transducers, and cur- 
rent transformers and shunts, current transducers. 

Another popular current-sensing device employs the 
Hall-effect principle. Hall-effect current transducers mea- 
sure the effect of an electromagnetic field on a semicon- 
ductor. Basically, such devices operate on the interaction of 
magnetic force and the movement of charge through a 
semiconductor. Consider figure 5.57 in which a current, I c , 
is flowing and a magnetic field is acting perpendicular to 
the current. The magnetic field will deflect the charge 
carriers in proportion to the field strength. This action 
produces a Hall-effect voltage, as shown, which is in 



127 



Vertical 

input 

terminals 

Vin f 



Attenuator, restricts 
the range of the 
signal to amplify 




Attenuator and 
preamplifier may be 
a combined plug-in 
unit for versatility 



Delays vertical deflection 
until sweep generator starts 
so full waveform is 
displayed on CRT screen 



Vertica 
amplifier 



External 

trigger 

input 



Internal 

trigger 

amplifier 



To internally synchronize sweep generator 
with the vertical signal 



Controls sweep 
generator for 
stable display 



o-o 



Vj 



Generates precise 
sawtooth voltage 
waveform with 
adjustable 
duration 



T 



Trigger 
generator 



Sweep 
qenerator 



r 



Attenuator 









Horizontal 
preamplifier 


o * » 










Horizonta 
amplifier 



Horizontal input terminals, 
used to drive horizontal 
deflection plates from 
external source if desired 



Attenuator and 
preamplifier may be 
a combined plug-in 
unit for versatility 




Electron gun 
controls, ace 
and focuses 
beam 



Vertical sea 
amplitude 

Horizontal 
scale, time 



Simplified block diagram 

Figure 5.56.— Cathode-ray tube. 



Sinusoid displayed with sweep 



Input 
resistance 



Magnetic 
flux, A 




Metallized 
contacts 



Hall voltage 



KEY 



x, y, z Dimensions of Hall-effect device 
Figure 5.57.— Semiconductor illustrating Hall effect. 

proportion to the magnetic field. Current flow in a conduc- 
tor produces an electromagnetic field that can be mea- 
sured by a Hall-effect device, thus producing a voltage 
output in proportion to the current. Most times, the 
magnetic field requires concentration. In some Hall-effect 
instruments, this is performed by a core of magnetic, 
low-retentivity material that can be clipped around a 
conductor. The semiconductor is mounted on the core and 



oriented at right angles to the induced magnetic field. The 
combined unit appears much like a split-core CT, and 
through this method dc as well as ac currents can be 
measured with high precision. Many instruments used for 
precise power-system measurements employ Hall-effect 
devices. 

The output from a transducer is sometimes incompat- 
ible with the input of the instrument, or for safety reasons 
is not isolated from the power system. In these instances, 
the signal requires conditioning, and electronic circuitry, 
usually amplifiers, is called upon to perform the task. 



INSTRUMENT INSTALLATIONS 

It is common to find several instruments included as 
part of major power-equipment circuitry. As a summary to 
this chapter, the following describes the typical locations 
for measuring instruments within a power system. 

1. The termination of utility transmission lines: 
Voltmeters, 
Ammeters, 
Wattmeters, 

Varmeters or power factor meters, 
Watthour meters, 
Demand meters, and 
Frequency meters. 



128 



2. Substation secondaries (outgoing distribution): 

Voltmeters, 

Ammeters, 

Wattmeters, 

Varmeters or power factor meters, 

Watthour meters (demand attachment optional), 

and 

lest blocks (or connection points) for portable 

instruments. 

3. Switchhouses, load centers, and rectifiers: 

Voltmeters, 

Ammeters, and 

Test blocks for portable instruments. 

4. Machinery: 

Voltmeters (optional), 

Ammeters, 

Elapsed-time meters (optional), and 

Watthour meters (also optional). 



Two points must be considered when applying the above 
listing. At the higher transmission voltages, say 69 kV 
and up, it is sometimes advantageous to have the utility 
metering point at the substation transformer secondary. 
This would eliminate item 1 and might add additional 
instruments to item 2 for reasons of economics. The capital 
required for high-voltage metering can prohibit its use. 
The test blocks listed for portable instruments are neces- 
sary because they provide an easy avenue for maintenance 
and trouble shooting. In general, the test blocks consist of 
a series of terminals to which permanent connections are 
made to important circuit portions, such as major compo- 
nents. The block is located on the surface of the equip- 
ment, accessible only to maintenance personnel, and can 
eliminate the need for some permanent instruments. 
However, as was mentioned at the beginning, voltmeters 
and ammeters are considered to be the minimum perma- 
nent instrumentation within mine power equipment. 



129 



CHAPTER 6.— MOTORS AND MOTOR CONTROL 



The subject of this chapter is the electromechanical 
conversion equipment that links electrical and mechani- 
cal systems and makes it possible to convert from one 
energy form to the other. The primary electromechanical 
devices are generators and motors. In generators, mechan- 
ical power is used to generate electrical power. Electric 
motors can be viewed as generators in reverse; they 
convert electrical power into mechanical power. The word 
"motor" can be applied to a device that converts energy of 
any form into mechanical power, but for purposes of this 
chapter the term is restricted to those machines that 
receive electrical energy. 

Generators have limited but important applications 
in most mining operations. The principal functions are in 
motor-generator (m-g) sets for surface excavating machin- 
ery and mine hoists, and for providing emergency power to 
ventilation fans and hoisting equipment. Motors, on the 
other hand, are used so extensively that they are the most 
important mechanical source in mining machinery and 
the most important loads on the mine electrical system. 
By far the majority of mines, milling plants, preparation 
plants, and other related mining activities would find it 
virtually impossible to operate without electric motors. 

Generators and motors can be studied independently 
of each other, but comprehension of motor operation is 
more easily obtained when generation is covered first. 
Consequently, the chapter will follow this format. In view 
of their relative importance, motors and their control will 
be the principal chapter discussion. The content will be 
elementary, but the objective is to provide sufficient infor- 
mation so that the effect of motors on the mine power 
system and specific motor applications in mining can be 
appreciated. 



it moves at right angles to the magnetic-field direction, a 
voltage will be induced in the conductor. Obviously, when 
a conductor is part of a closed loop, current will flow. The 
direction of the current produced depends upon the direc- 
tions of the magnetic flux and the conductor movement. 
Fleming's "right-hand rule" is often used to determine the 
current direction, and it also helps to illustrate the inter- 
relationships among the three dependent parameters. The 
thumb, forefinger, and center finger of the right hand are 
stretched out so they are mutually at right angles to each 
other. If the hand is placed such that the forefinger points 
in the flux direction, with the thumb pointing in the 
direction of conductor motion, the center finger points in 
the direction of current flow. 

When the conductor cuts a magnetic field at a specific 
relative velocity, it has been found (3) 1 that the instanta- 
neous magnitude of the voltage induced, e, can be calcu- 
lated by 



e = Bfv„ 



(6.1) 



where B = magnetic field flux density, T, 

2 = conductor length, M, 
and v c = conductor velocity at right angles to magnetic 
flux field, m/s. 

Accordingly, the magnitude of induced potential de- 
pends upon the flux density, the conductor length, and the 
conductor velocity relative to the magnetic field. By vary- 
ing these parameters, a voltage of almost any magnitude 
can be theoretically produced. It is this electromechanical 
principle, an adaptation of Faraday's law, that is utilized 
in the generation of alternating and direct currents (ac 
and dc). 



ALTERNATING CURRENT GENERATION 

In chapter 2, it was demonstrated that a voltage is 
induced in a conductor when there is relative motion 
between the conductor and a magnetic field. This electro- 
magnetic induction concept, called Faraday's law, is the 
basic principle behind the generation of voltage in electric 
machines. The following paragraphs return to this funda- 
mental concept but in a slightly different fashion and 
serve as a transition between induced voltages in induc- 
tors or transformers and generators. 

In figure 6.1, a conductor is under the influence of a 
magnetic field. If a force is placed on the conductor so that 



Force on conductor 
producing V c 




Induced emf, e 



Figure 6.1. —Production of voltage from magnetic field 
(emf = electromotive force). 



Principle of Generator Operation 

Figure 6.2 illustrates the basic principle of ac gener- 
ation. Consider a loop of conductor mounted on a mechan- 
ical drive shaft through an insulating block and rotated in 
a magnetic field (9). A circular metallic ring, called a slip 
ring, is connected to each end of the loop, and brushes 
contact each slip ring surface to allow the connection of 
stationary conductors. 

When the loop is rotated by a mechanical drive, a 
potential is induced in each side of the loop that is 
proportional to the conductor velocity at right angles to 
the magnetic field. At the position shown in figure 6.2, the 
induced voltage is at a maximum (relative right-angle 
velocity is maximum), but at 90° from this position there 
will be no induction (relative velocity is zero). The instan- 
taneous voltages for the entire loop can be added algebra- 
ically as they are in series. Continued rotation will pro- 
duce a sinusoidal voltage (fig. 6.2.B) and thus sinusoidal 
current, if the loop is part of a closed circuit. 

Generator Construction 

It has been shown that there are basically two re- 
quirements for generation. The first is a conductor or 



1 Italicized numbers in parentheses refer to items in the list of references 
at the end of this chapter. 



130 



Slip ring- 




Axis of stator field 




Vertical axis 



B 



Figure 6.2.— Demonstration of ac generation. 



winding in which a desired voltage is to be induced. This 
is termed an armature winding, and the structure enclos- 
ing it is called an armature. The second requirement is a 
magnetic-field source, and this is normally created by a 
field winding, although some very small machines use 
permanent magnets. To classify the rotating and fixed 
machine portions, the rotating member is referred to as 
the rotor, while the stationary portion is the stator. 

The windings are placed on two concentric cylindrical 
iron cores with a small air gap between so that the flux 
path in the machine is as efficient as possible. The inner 
core usually serves as the rotor. Thin laminations, at times 
insulated from each other, are employed to minimize 
eddy-current loss, as in transformer construction. The 
structure of either core is one of two types: salient poles or 
nonsalient poles, which form the center lines of the 
magnetic field. Salient poles stick out from the cylinder 
surface (fig. 6.3) and have the windings around them but 
located near the core surface in the vicinity of the air gap. 
Nonsalient poles are part of a completely cylindrical 
surface, with the windings positioned in slots (fig. 6.4) (3). 
The conductors are usually insulated from the core. Sur- 
rounding the cores and windings is a structure called the 
frame, with some form of end enclosure. The frame serves 
to anchor the stationary machine elements to a founda- 
tion. The end enclosure may contain bearings, of either 
sleeve, ball, roller, or needle types, that support the rotor 
shaft and position the rotor properly with respect to the 
stator. Figure 6.5 is a sketch illustrating these physical 
components. 

The function of an electromechanical machine is 
commonly described in terms of the number of available 
magnetic poles in the field. Thus, the elementary machine 



Stator 




Axis of rotor field 



Figure 6.3.— Cross section of machine with salient poles on 
stator and nonsalient poles on rotor. 



r^— Axis of stator field 




Axis of rotor field 



Air gap 



Figure 6.4.— Cross section of machine with nonsalient poles 
on stator and rotor. 




Figure 6.5.— Simplified sketch of electromechanical 
machine illustrating physical components. 



in figure 6.2 is a two-pole generator. Most generators, 
however, have more than two poles, usually even numbers 
of four, six, eight, and so on. Figure 6.6 shows an elemen- 
tary four-pole generator (9). Here the armature needs only 
to turn 180° to produce a full sinusoidal cycle in its 
winding output. 



131 



The foregoing terminology applies to all electrome- 
chanical rotating machinery. For practically all ac gener- 
ators or alternators, the armature is contained in the 
stator. The field winding is part of the rotor, with a dc field 
current supplied through slip rings, which is the reverse of 
the situation discussed previously. 

Three-Phase Generation 

A preliminary discussion of three-phase power gener- 
ation has already been presented in chapter 4 but only in 
the context of balanced three-phase systems. This section 
will elaborate on its electromechanical conversion. 

Consider figure 6.7, a cross-sectional view of an ele- 
mentary three-phase, two-pole generator. The machine is 
termed two-pole because of the number of available mag- 
netic poles in the field winding. Located in the stator, the 
armature has three single-conductor windings a, b, and c, 
whose axes are 120° apart. The rotor containing the field 
winding is turned at a constant speed by a mechanical 
power source connected to the rotor shaft, and the field 
winding is excited by dc. The magnetic-flux distribution 
around the air-gap circumference of the machine is de- 
signed so it forms a sine wave (3). Therefore, the induced 
voltage in each armature winding varies sinusoidally with 
the familiar 120° displacement among the three gener- 
ated potentials. 

For this two-pole generator, the sinusoidal voltage in 
each phase winding goes through one full cycle per rotor 
rotation. The waveform frequency (hertz) is identical to the 
rotor speed (revolutions per second). The sinusoid is thus 
in time with or synchronized with the mechanical speed, 
and such ac generators are often termed synchronous 
generators. For 60-Hz output, rotor speed is 60 r/s or 3,600 
r/min. 

An example of an elementary four-pole generator is 
shown in figure 6.8A. Here, the rotor poles alternate 
between north or south polarity when rotated. Each phase 
of the armature consists of two windings connected in 
series, as shown in figure 6.8S. The induced voltage per 
phase thus completes two cycles for each rotor revolution. 



DIRECT CURRENT GENERATORS 

A very elementary two-pole dc generator is shown in 
figure 6.9. The illustration differs from figure 6.2 only in 
that the dc generator has a commutator in place of the slip 
rings. The commutator is an annular ring that is split into 
parts (in this case, two), which are insulated from each 



dc 
for exciting 



Output 




Figure 6.6.— Elementary four-pole, single-phase ac 
generator. 



Armature structure 
or stator 



Field structure 
or rotor 




Armature winding 



Field winding 

excited by dc 

through slip rings 



Field pole produced 
by dc in field 
winding 



Figure 6.7.— Elementary two-pole, three-phase generator. 





A Cross section B Armature schematic 

Figure 6.8.— Elementary four-pole, three-phase generator. 





01 2345670 
O 45° 135° 225° 315° 
90° 180° 270° 

Generator Voltage output versus time 

Figure 6.9.— Demonstration of dc generation. 



132 



other. Each part is termed a commutator segment. As 
before, carbon brushes contact the ring surface to allow 
connection of stationary conductors. During armature 
rotation, the voltage produced in the loop is a sinusoid, as 
in the ac generator. The commutator serves to rectify the 
waveform mechanically since at all times the positive and 
negative brushes are connected with the correct armature- 
winding polarity. In other words, the connection to the 
loop reverses or commutates every one-half revolution. 
Thus, the generator output waveform is the same as 
full-wave rectification (fig. 5.5). If the rotational direction 
reverses, so does the brush polarity. 

Because of the ripple voltage, the two-pole dc genera- 
tor is not realistic. In practical dc generators, the arma- 
ture consists of many windings, with the commutator 
having a corresponding number of segments (fig. 6.10). In 
such a case, current from the generator will never drop to 
zero. When the number of armature windings is increased, 
the output ripple voltage decreases, and the average direct 
voltage will be closer to the peak voltage. 

Unlike ac generators, dc generators have the arma- 
ture winding on the rotor and the field winding in the 
stator. The field must be excited by dc provided by a source, 
which may be either external or internal. The internal 
excitation is possible because the armature is a dc source 
and can supply current to the field as well as the load. 
However, in order to start generation, the stator core of 
these machines must have residual magnetism. Gener- 
ators connected in this way are called self-excited. When 
the source is external, the generator is termed separately 
excited. This is diagrammed in figure 6.11. 

Self-excited generators have three configurations, de- 
pending on the field-winding connection: series, shunt, 
and compound, as shown respectively in figures 6.12, 6.13, 
and 6.14 The terms series and shunt relate directly to the 
winding connections. The compound generator has two 
windings, one connected in series and the other shunting 
the armature. 

Each of the generator connections has a characteristic 
voltage output versus load current (3). Because the field, 
armature, and load currents are the same in series gener- 
ators, the output voltage fluctuates widely with the load. 




Brush 

Commutator 
Brush 



Generator 




90° 180° 270° 

Voltage output versus time 



360 c 



Figure 6.10.— Dc generator with two armature windings at 
right angles. 



Hence, this connection is rarely used. Although shunt 
generator voltage output drops slightly as load current is 
increased, the regulation is satisfactory for many pur- 
poses. Compound generators are normally connected so 
the magnetic actions of the shunt and series windings aid 
each other. The resultant magnetic flux of the field can 
increase with load current, causing the output voltage to 
remain nearly constant. The level of output voltage in both 
the shunt and compound generators can be controlled by 
the variable resistance in series with the shunt field 
winding. The resistance in the separately excited genera- 
tor provides the same function, but precise output-voltage 
control is obtained because the field-winding current is not 
a function of the load. 



Field winding 



To dc source 




To load 



Armature 



•Brush 
Field rheostat 



Figure 6.11.— Separately excited dc generator. 



Field winding 



Armature 
Brush 




Figure 6.12.— Series dc generator. 




Field rheostat 



To load 



■Field winding 
Figure 6.13.— Shunt dc generator. 



Armature - 
Brush 



/-Series field 
/ /-Field rheostat 



P 



To load 



Shunt field 
Figure 6.14.— Compound dc generator. 



133 



MOTOR BASICS 

The essential motor parts are similar to those of a 
generator and include 

• Two concentric cylindrical laminated-iron cores, 
separated by an air gap, to carry magnetic flux; 

• Two sets of windings, wound or embedded in slots in 
the iron cores, either or both excited by dc or ac; and 

• The inactive motor elements, including the frame, 
end bells, bearings, and so forth. 

Combinations of these parts are found in practically all 
motors. Motors employ electrical energy to produce me- 
chanical force, which is the reverse process from generator 
operation. The force of interest in motors is that which 
tends to produce rotation, or torque. 

Torque 

Motor torque considerations are based on the funda- 
mental principle that a mechanical force is exerted on a 
current-carrying conductor in a magnetic field. A graphic 
example of this situation is shown in figure 6.15. Here, the 
magnetic field that surrounds the conductor (due to its 
current) interlinks with the largest magnetic field. This 
creates a large concentration of magnetic flux at one side 
of the conductor, which tends to force the conductor toward 
the lesser flux concentration. The result is an instanta- 
neous force, f, at right angles to the magnetic field. 

The magnitude of the force depends upon the mag- 
netic field flux density, the conductor length, and the level 
of instantaneous current, and can be calculated for a 
straight conductor by (3) 

f = Bti, (6.2) 

where f = force, N, 

B = magnetic field flux density, T, 

t = conductor length, m, 
and i = instantaneous current, A. 

If the conductor is fixed by a radial distance, r, from the 
center of a rotor shaft, the associated torque, T, is (3) 



T = Blri, 

where T = torque, N-m, 

and r = radial distance or moment arm, m. 



(6.3) 



For a winding, the total torque is the summation of the 
torques for the individual conductors or coil sides. For 



electromechanical machines, this mechanical quantity is 
termed electromagnetic torque, and when combined with 
rotation the resultant power quantities follow the rules of 
mechanics. 

Another way of visualizing the development of motor 
torque is the interaction of two magnetic fields. A mechan- 
ical force is exerted on magnetic material, be it a perma- 
nent magnet or magnetism created by electric current 
flow. The force tends to align the material with the closest 
part of a magnetic field, so the north pole of one machine 
member is directly in line with the south pole of the other 
member. If the force is acting at a moment arm about a 
rotor shaft, torque is produced. 

Even though equation 6.3 is expressed in newton 
meters (adhering to SI units), the quantities normally 
used are pound-feet, ounce-inches, and gram-centimeters. 
The common method of relating motor mechanics is by 
reference to a percent of full-load torque. 

Speed-Torque Relationships 

Speed-torque curves are the mechanical characteristic 
curves of a motor; a general example for an induction 
motor is provided in figure 6.16 for discussion (15). One 
application for these curves is to find the most suitable 
drive for a given machine. As the machine load can also be 
described by a speed-torque curve (see load torque in the 
figure), the comparison of the load and motor curves will 
show if the motor has the necessary characteristics to 
drive the load and also what the operating point will be. 
The operating point is the intersection of the two curves. 
Many other parts of the motor characteristic relate its 
suitability for a specific application, and some of these are 
listed below and are shown in the labels of figure 6.16. 

1. Lockedrrotor torque. The minimum torque devel- 
oped by a motor at the instant of power application, 
sometimes called breakaway or starting torque. 

2. Accelerating torque. The torque developed during 
the period from zero to full rated speed with rated power 
applied. The term is often used for the net torque between 
the motor and the load. It is apparent in the figure that 
this is a nonlinear value with speed. 

3. Breakdown torque. The maximum torque possible 
from the motor with rated power input, also called maxi- 
mum torque. 

4. Pullup torque. The minimum torque developed 
during motor acceleration from zero to full rated speed 
with rated power applied. The minimum can exist in some 
motors at full rated speed. 

5. Full-load torque. The torque necessary to provide 
rated output at rated speed with rated power applied. 




Conductor carrying current 
in a magnetic field 




Flux produced by 
conductor with 
respect to field 



Force on 
conductor 




Resultant distortion 
of magnetic field 



Figure 6.15.— Current-carrying conductor in a magnetic field. 



134 



6. Pullout torque. The maximum torque produced by a 
motor without stalling. This is sometimes incorrectly 
referred to as the maximum or breakdown torque. If a 
torque is applied to a motor above this value during 
operation, it will stall. 

The other motor terms listed in figure 6.16 are tied to 
specific motor types or operations yet to be discussed. 

Standardization 

The National Electrical Manufacturers Association 
(NEMA) sets standards for the manufacture of electric 
motors (15), which are used throughout the mining indus- 
try. NEMA standards generally cover seven areas: speed- 
torque characteristics, frame size, enclosure, horsepower 
rating, voltage, temperature rise, and application. Al- 
though machines from different manufacturers should be 




Full-load torque 
Synchronous 
torque 



Zero Speed Synchronous 

Figure 6.16.— General speed-torque motor characteristic. 



directly interchangeable when they conform to a particu- 
lar NEMA standard, there may still be some variation 
between manufacturers. 

Frame Size 

Most motors of 250 hp and under are rated according 
to a frame number that specifies the essential mounting 
dimensions (fig. 6.17) (15). The same frame number series 
covers all ac or dc motor types, and a dozen or more 
different motors might have the same frame. 

Enclosure 

Motor enclosures are usually classified as open or 
totally enclosed. Open motors simply have openings, usu- 
ally in the end plates, to allow air cooling of the windings. 
Totally enclosed motors prevent passage of air into the 
enclosure, but these motors are not always sufficiently 
closed to be air tight. In this general class are the 
explosion-proof motors (see chapter 16), dust-ignition-proof 
motors, dust-tight motors, and waterproof motors. Cooling 
for these motors can be by air conduction on the outer 
frame, internal forced air through a pipe, or a liquid-cooled 
(water or oil) outer jacket surrounding the frame. 

Horsepower 

Motor horsepower is also standardized, such as 1/2, 
3/4, 1, 1-1/2, 2, 3, 5, 7-1/2, 10, 15, 20, 25, 30, 40, 50, 60, 75, 
100, 125, 150, 200, and 250 at speeds for 2 to 16 poles with 
60-Hz operation (15). Above 250 hp, the standard powers 
are related to motor type. When a horsepower is given, it 
is often combined with a service factor to allow for usual 
fluctuations in supply voltage or slight overloading. The 
service factor indicates the permissible overload and is a 
multiplier applied to the normal horsepower rating, with 
values ranging from 1.0 to 1.4 depending on the size and 
type of motor. For instance, a 1.15 service factor would 



W-*.SI 



w — w 



BA 

H-diam, 
4 holes 



■N minus W 



■XE 



t 

D 

1 


\ 1 


r J 


i 


<! 


— E-^r 

A M 



XD-*| 




SECTION AA 



No. 


A 


B 


D 


E 


F 


BA 


H 


NW 


U 


V 
mm 


XC 


XO 


XE 


I45T 
I84T 
2I3T 


7 

9 

10-1/2 


6 
7-1/2 
7-1/2 


3-1/2 
4-1/2 
5-1/4 


2-3/4 
3-3/4 
4-1/4 


2-1/2 
2-3/4 
2-3/4 


2-1/4 
2-3/4 
3-1/2 


11/32 
13/32 
13/32 


2-1/4 

2-3/4 

3-3/8 


7/8 

1-3/8 

1-3/8 


2 
2-1/2 
3-1/8 


3/16 

1/4 

5/16 


3/16 

1/4 

5/16 


1-3/8 
1-3/4 
2-3/8 



Dimensions in inches 
Figure 6.17.— Examples of three frame number dimensions. 



135 



indicate that a motor can carry 15% more than rated load 
continuously without overheating (that is, exceeding the 
rated temperature rise) as long as the frequency, ambient 
temperature, line voltage, and so on are at rated values. 
An interesting situation occurs with frame sizes pri- 
marily intended for motors of less than 1 hp {fractional 
horsepower). A fractional-horsepower motor is considered 
to be any motor built in a fractional-horsepower frame, 
even if the actual horsepower rating is in excess of 1 hp. 
Most of these motors, however, are only available for 
single-phase ac. 

Voltage 

NEMA voltage designations are specified for most 
motors, be they three-phase, single-phase, or dc (15). A 
listing of the voltage ratings (also called the motor termi- 
nal voltage) common to mining is available in table 6.1. A 
maximum voltage variation of + 10% from rated is per- 
mitted. With ac motors, the allowable frequency fluctua- 
tion for the power supply is + 5%. 



Table 6.1. — Motor voltage ratings common to mining 



System type n 



3-phase, low voltage . 



3-phase, medium voltage. 
3-phase, high voltage 



Direct current. 
Single phase . 



Nominal system 
voltage, V 



Motor rated 
voltage, V 



2 208 


208 


230 


220 


3 460 


3 440 


"480 


4 440 


3 575 


3 550 


4 600 


"550 


4 1,040 


4 950 


2,400 


2,300 


4,160 


4,000 


7,200 


6,600 


12,470 


12,470 


13,200 


13,200 


13,800 


13,200 


300 


250 


600 


550 


120 


115 


240 


230 



System voltage designations follow 30 CFR 18, 75, 77. 
: Wye. 

' More suitable for stationary equipment applications. 
1 Intended use is mobile mining equipment. 



Temperature Rise 

The allowable temperature rise from an ambient 
temperature is dependent upon the class of insulation 
used in the motor (15). The electrical insulation system is 
one of the most important components of a motor, as its 
degradation seriously affects the reliability and service 
life of the motor. Insulation systems are divided into four 
classes, A, B, F, and H, depending upon their thermal 
endurance. Ratings for industrial motors are typically 
based on a 40° C ambient temperature, but some are 
based on 25° C. Table 6.2 provides the allowable rise for 
each insulation class; also included are the common insu- 
lating materials and the maximum "hot-spot" tempera- 
ture, which is the highest temperature allowable at any 
part of the motor. Motors built to a specific class and 
operated so the recommended temperatures are not ex- 
ceeded may be expected to have a serviceable life of 20 yr 
with minimal maintenance. However, physical abuse and 
the electrical stresses discussed in chapter 11 can seri- 
ously shorten the motor life regardless of insulation class 
and operation. Class A insulated motors are rare in almost 
all applications, except for very small horsepowers. Class 
F and H insulations are most often used in motors for 
mining applications (17). The maximum surface tempera- 
ture of any permissible mining motor must not exceed 
150° C (see chapter 16). 

If the maximum ambient temperature is greater than 
specified, the allowable temperature rise must be de- 
creased by the difference in temperature above ambient. 
Maximum ambient is the highest temperature the motor 
is normally exposed to. The rise may be increased by a like 
amount when the maximum ambient temperature is be- 
low that specified. These specifications are applicable 
when operating under typical barometric pressure, as long 
as the altitude does not exceed 1,000 m (3,300 ft). Above 
1,000 m, the allowable temperature rise must be reduced 
1.0% for each 100 m (330 ft) above 1,000 m. 

Additional classification standards will be discussed 
in the following sections, for example, applications as to 
load-speed and load-torque requirements. These are tied to 
the torque-speed characteristics, which are related to the 
motor type. 



Table 6.2.— Motor insulation classes 



Temp rise, °C Maximum 





nsuia- 
tion 


Open 


Totally 


hot-spot 


Common insulating 




class 


motors 


enclosed 
motors 


temp, 
°C 


materials 2 


A 




50 


55 


105 


Cotton, cellulose, paper, 
organic, enamel-coated wire. 


R 




70 


75 


130 


Mica, glass fiber, asbestos. 


F 




90 


95 


155 


Do. 


H 




105 


115 


180 


Mica, glass fiber, silicone 
elastomers, silicone resins, 
asbestos. 



1 Allowable rise from ambient temperature. 

2 Each class has compatible bonding agents for the materials shown. 



Motor Type 

The classification of motor types, which is dependent 
upon how the stator or rotor windings are excited, results 
in three general motor classes: induction, synchronous, 
and dc. The first two are ac machines, and for many 
applications these are more rugged, require less mainte- 
nance, and are less expensive than dc motors of equal 
horsepower and speed ratings. Ac motors can be used 
effectively for the majority of motor applications except 
when very high starting torques are required. The most 
widely used ac type is the squirrel-cage induction motor, 
so-called for its appearance. It has no slip rings, commu- 
tator, or brushes to wear out and uses the simplest kind of 
starting equipment. Three-phase squirrel-cage induction 
motors and series-wound dc motors are the most popular 
electromechanical machines in mining. 

After the presentation in chapters 2, 3, and 4, it could 
be expected that three-phase motors would be more com- 
plex than their single-phase and dc counterparts and 
therefore more difficult to understand. However, although 
some parts of three-phase motor construction are more 
complex, their operation is simpler. As just stated, induc- 
tion and synchronous machines are the two major ac motor 
types. Synchronous motors correspond to three-phase gen- 
erators. In typical large machines, dc is applied to the field 
winding located in the rotor, while three-phase ac (instead 



136 



of being generated) is supplied to armature windings 
placed in the stator. Induction motors also receive ac power 
at the stator windings, but ac is delivered to the rotor 
winding indirectly by induction, in the same manner as in 
a transformer. 



THREE-PHASE SQUIRREL-CAGE INDUCTION 
MOTORS 

In order to comprehend the operation of induction 
motors and understand important terminology, it is per- 
haps best to start with a simple demonstration. Although 
the motor does not have familiar motor components, its 
construction and operation do have direct application in 
induction-disk relays and watthour meters. 

Consider figure 6.18, which depicts an aluminum disk 
and a horseshoe magnet (8). Both are mounted about the 
same axis and are free to rotate. When the magnet is 
rotated, the disk cuts the magnetic lines of force, a voltage 
is induced, and eddy currents will then flow. Under the 
magnet's south pole, the eddy currents set up north 
magnetic poles, and conversely for the north pole of the 
magnet. Because the pole attracts, the disk rotates, follow- 
ing the magnet. The disk can never reach the magnet 
speed, as there would be no relative motion between the 
two (that is, no induction would exist). The mandatory 
difference in speed for induction motors is called slip. 

In conventional induction motors, the action of the 
disk occurs in a rotor winding, and the rotating magnetic 
field is supplied by the stator winding. For induction-disk 
relays, the aluminum disk is the same but the induction 
force is supplied by an ac-driven stator (see chapter 9 and 
also the description of watthour meters in chapter 5). 

Elementary Three-Phase Motor 

Figure 6.19 illustrates an elementary two-pole, three- 
phase squirrel-cage induction motor (11). The stator con- 
sists of three salient poles spaced 120° apart; the stator 
windings around each pole are connected in wye and 
energized by a three-phase system. The rotor has three 
main elements: a shaft (not shown), core, and winding. 

As with generators, the rotor core is made of iron 
laminations pressed onto the shaft. The squirrel-cage 
winding is constructed by embedding heavy copper or 
aluminum bars in the core slots. The bars are connected to 
each other by copper or aluminum rings located on both 
core ends, which complete the closed circuit. In other 
words, there are no external connections to this rotor 
winding either by slip rings or a commutator. Figure 6.20 
shows the winding construction. 

When the stator windings are powered by a three- 
phase system, currents through the coils reach their 
respective maxima at different intervals in time. Since the 
three currents are displaced by 120°, the magnetic field 
generated by each coil is also displaced from the other two 
by 120° (fig. 6.21A). The magnetic field of each winding 
alternates from north to south; thus, each has the action of 
two poles. Figure 6.21B shows the instantaneous direction 
of a stator flux as it passes through the rotor at different 
time intervals. At zero degrees, for instance, phase A is at 
maximum north, while phases B and C are weak south 
poles. At 60°, phase C becomes strongly polarized in the 
south direction and phases B and A are weak norths. The 
larger arrows shown in the figure represent the instanta- 



String 




Permanent 
magnet 



Copper or 
aluminum 




Direction of 

induced 
eddy currents 



Rotation 



^lurntabie \tf magnet/V "' <» ■'• 

Iron plate 

B Top view 
Figure 6.18.— Demonstration of induction-motor operation. 



Bearing-^ -v_pj vot 
A Front view 




Figure 6.19.— Elementary three-phase induction motor. 




Figure 6.20.— Squirrel-cage rotor winding. 



neous direction of the resultant two-pole magnetic field. 
Consequently, for this example, the magnetic field is 
rotating counterclockwise. 

In a transformer, voltages are induced in the second- 
ary circuit by the primary. The stator of an induction 
motor acts in the same manner as the primary, with the 
rotor winding acting as the secondary winding. The rotat- 
ing magnetic field of the stator cuts the rotor conductors, 
and motor action is developed. The relative motion, or slip, 
between the rotating flux and the rotor generates voltages 
within the rotor conductors. 

According to Lenz's law, the voltage induced in each 
rotor bar will be in a direction opposing the relative 



137 



motion of the rotating flux and the rotor. The induced- 
voltage direction in the rotor conductors under the influ- 
ence of a two-pole rotating magnetic field is shown in 
figure 6.22, where positive implies that the voltage direc- 
tion is toward the viewer. The voltage magnitude in each 
bar depends upon the stator magnetic-field density at that 
point. These voltages cause currents to flow through the 
bars, an end ring, adjacent bars, and then back through 
the other end ring to the origins, in complete loops. The 
circulating rotor currents produce magnetic fields about 
each rotor bar. 

The interaction between the stator field and the fields 
around the rotor conductors results in a mechanical couple 
and thus motor torque. Hence, the rotor will rotate in the 
same direction as the stator field. A simple reversal of any 
two phase conductors to the stator windings of a three- 
phase induction motor will reverse the stator phase se- 
quence and thus reverse the motor rotation. 



The speed at which the stator field rotates is termed 
the synchronous speed of the motor and can be calculated 
from 



120f 



n Q = 



(6.4) 



where n s = synchronous motor speed, r/min, 

f = line frequency, Hz, 
and p = number of magnetic poles presented by stator. 

As slip is required to produce rotor induction, a squirrel- 
cage motor may approach but never obtain synchronous 
speed. Slip can be expressed mathematically as 



s = 



gs-nr 



(6.5) 



ABC 




Electrical degrees 




& j&S&Q. &&>$. B %V^# c 

0° 60° 120° 180° 




240° 



300° 



360° 



Figure 6.21.— Rotating magnetic field in elementary three- 
phase, two-pole induction motor. 



where s = motor slip, expressed as a per-unit decimal or a 

percent, 
and n,. = actual motor speed, r/min. 

Losses will occur in actual motors because of electrical 
and mechanical inefficiencies. Those prominent in induc- 
tion motors are 

• Rotor winding loss, related to I 2 R; 

• Stator winding loss, also an I 2 R loss; 

• Stator core loss, caused by eddy currents and hys- 
teresis in the core iron; and 

• Friction and windage (rotational or mechanical) 
losses. 

These are almost pure active powers; therefore they are 
often expressed in watts. The losses in both windings of 
induction motors vary as the square of line current, core 
loss is nearly constant, and unless motor speed varies 
considerably rotational losses are nearly constant (11). 
Knowledge of machine losses allows the determination of 
motor heating and efficiency. 

The efficiency of motor operation is a measure of the 
ability to convert input power to mechanical power: 



Efficiency = 



output input - total losses 



input 



input 



(6.6) 



which may be expressed as a per-unit decimal or a percent. 
Slip is also related to motor efficiency, being numerically 
equal to the ratio of winding loss in the rotor to the total 
rotor power input: 



Rotor conductor 




Figure 6.22.— Induced rotor potential by rotating flux. 



s = 



rotor winding loss 
rotor power input 



or 



s = 



rotor winding loss 



motor power input - stator losses ' 



(6.7a) 



(6.76) 



Stator losses in equation 6.76 include friction and wind- 
age. Equations 6.4 through 6.7 can be employed to calcu- 
late the synchronous and actual motor speeds and also the 
possible power and torque output, realizing that (8) 

power output (watts) = 746 (horsepower output), (6.8a) 



138 



and 



in which 



hp = 



5250 



T = 



KI^Rr 



(6.86) 



where hp = horsepower output of motor, 
n,. = actual motor speed, r/min, 
T = motor torque, ft-lb, 
K = a torque constant, ft-lb/V, 
I,. = rotor current, A, 

and R,. = rotor resistance, Q. 



^ ^ ^ 

C B A 




Figure 6.23.— Lapped windings of three-phase motor stator. 



Motor Construction 



The elementary salient-pole motor of figure 6.19 is 
undesirable from the standpoint of the ineffective use of 
material and space, as well as its overall inefficiency. The 
main disadvantage is coupled to the distinguishable stator 
poles. To overcome this problem, actual induction motors 
have lapped stator coils, as shown in figure 6.23, where 
several coils make up a stator winding that can be either 
delta or wye connected. The flux directions of each coil are 
illustrated as <j> c , </> B , 4> A , and each coil contributes to the 
rotating flux development of the entire stator. The coils 
and windings are arranged to have the same effect as 
salient poles, but the poles are not physically distinguish- 
able. An induction motor is assigned a specific pole num- 
ber if at any given instant the stator windings set up the 
same number of magnetic pole fields. 

The rotor core and squirrel-cage conductors are usu- 
ally not insulated from each other, because the induced 
current is effectively contained within the conductors 
owing to their significantly lower resistance. The rotor 
core is pulled magnetically toward the stator core across 
the air gap. If the force is uneven when the rotor turns, the 
result is vibration. This is detrimental in several ways as 
it can lead to structural insulation failures, premature 
bearing failures, and misalignments with the motor load. 
Vibration does not occur if the magnetic effect about the 
rotor periphery is equal. An additional method for pre- 
venting vibration is to place rotor conductors in slots 
skewed to the stator slots so that a rotor slot passes 
gradually under a stator slot rather than abruptly. This 
practice also prevents "dead spots," or positions of near- 
zero or minimum magnetic influence. Another method of 
eliminating dead spots is to construct the motor so that 
the number of rotor slots plus the stator slots sums to a 
prime number. 

Motor Behavior 

Figure 6.24 is a graph of the speed, efficiency, power 
factor, power input, and current load of a typical three- 
phase induction motor found in mining applications. Fig- 
ure 6.25 shows a representative torque-speed characteris- 
tic for a similar machine. These curves can be used to 
describe the electrical and mechanical operation of induc- 
tion motors under loading. 

From the typical torque-speed curve, the torque at 
locked rotor is approximately 150% of rated. The level 
increases steadily with rotor acceleration to the maximum 
or breakdown torque. With applied power input, the rotor 
continues to accelerate until the slip reduction reduces the 



5 

(£* 
UJ 

5 < 

t- LJ 
=> £E 

15 

O 

H 

o 

2 



180 i- 



150 - uj 



120 - ft 



90 - 



60 



30 





1 1 ■ t— L.^ 
rpm 


— i— I — i — 1—1 — 




eff 


j 


90 


- ^^^pT 




80 
70 


/ X ° 

I / ° 




60 
50 


1 / *3 

J LL. 




40 






30 


A / s 




20 
10 


1 I I I 


i i i i i 



1,200 
1,150 
1.100 



10 20 30 40 50 60 70 80 90 I00 IIO 
OUTPUT, hp 



Figure 6.24.— Characteristic curves of three-phase induc- 
tion motor. 




f Rated torque 



50 1 00 
SYNCHRONOUS SPEED. % 
I I 



1 00 



50 
SLIP. % 



Figure 6.25.— Typical torque-speed characteristic for 
general-purpose induction motor. 



rotor current to a point where torque is equal to the load 
torque. 

Consider the motor running with no load. As the 
motor is loaded, slip increases, causing an increase of 
induction in the rotor. Hence, rotor current rises, resulting 
in a stronger rotor magnetic field and motor torque. 
Torque continues to increase with the increased shaft load 



139 



until breakdown torque is reached. Any further load 
results in a slip value that decreases torque. If the high 
load is sustained, the rotor will stop. 

Because the induction motor operates basically as a 
transformer, its electrical characteristics, as seen by the 
power source, will be a reflection of those occurring in the 
stator winding. Figure 6.26 shows phasor diagrams for 
rotor current and voltage during three operation points; 
these are referenced to the flux-density phasor of the 
stator, B stator (11). 

The rotor bars are embedded in the steel core so they 
have a high reactance (3). At locked-rotor conditions (rotor 
stationary), the stator magnetic field rotates past the 
motor at synchronous speed, and the induced voltage in 
the rotor conductors has the same frequency as the stator 
(or line frequency). The result is a high ratio of rotor 
reactance to resistance, and stator current lags stator 
voltage by a large amount (fig. 6.26A). During rotor 
acceleration, slip decreases, which also lowers the fre- 
quency of rotor current and voltage according to the 
following relationship (8): 



The torque developed by a three-phase induction mo- 
tor varies as the square of the stator supply voltage, or 



f r = sf, 



(6.9) 



where f r = frequency of sinusoidal voltage and current 
induced in rotor bars, Hz, 
s = slip, expressed as a decimal, 
and f = frequency of voltage and current in stator, Hz. 

Thus, inductive reactance drops, increasing the power 
factor (fig. 6.265). Theoretically, if the motor could obtain 
synchronous speed, the rotor power factor would reach 
unity (fig. 6.26Q. However, as this cannot happen in 
actual squirrel-cage motors, the maximum power factor is 
seldom greater than 0.85 (fig. 6.24) and never greater than 
0.95. 

Because the output torque increases with slip, motor 
speed decreases slightly as the load increases from no load 
to full load. Yet efficiency and power factor drop rapidly on 
low load conditions. Hence, an induction motor should not 
be operated at much below rated load for any length of 
time. It is apparent from figure 6.24 that efficiency dimin- 
ishes when motor load increases above a given value. 
Consequently, an induction motor should not be over- 
loaded for any extended period. Power-factor and efficiency 
curves normally follow roughly the same path; thus, power 
factor can be considered as an estimate of motor operating 
efficiency. 



Torque oc V^ 



(6.10) 



Therefore, a 10% reduction from rated stator voltage will 
cause a 19% reduction in available torque output. 

Insulation 

Insulation in motors normally has five forms: strand, 
turn, lead, crossover, and ground (15). Since the rotor 
conductors are uninsulated, the insulation of the stator 
winding conductors is the critical concern. The primary 
insulating system is that between the windings and the 
stator core or ground, and the secondary insulation is in 
strands, turns, leads, and crossovers. 

Copper magnet wire, and to a much lesser extent 
aluminum magnet wire, is used to construct the stator 
winding or coils. Strand insulation is most frequently a 
resinous coating on the wire. Turn insulation is applied 
after strands are wound into coils (or the actual windings), 
and this may be a resinous coating, resinous-film taping, 
paper taping, or a fibrous wrapping. These types of turn 
insulation are utilized for applications of 6,600 V and less; 
for higher voltages, additional layers of mica or varnished 
cloth tape can be used. Crossover insulation is employed to 
protect wires that cross each other. The crossovers are 
often the weakest point in winding construction; thus, 
they require additional protection. Lead insulation is 
simply insulation about the conductors leading to the 
windings. Lastly, ground or ground-wall insulation is the 
major insulation system of the motor and isolates the 
windings from the core. This insulation is always sub- 
jected to the highest potential difference and requires the 
most attention. 

Design Characteristics 

Figure 6.27 illustrates the standard NEMA torque- 
speed characteristics for squirrel-cage induction motors. 
The shapes of these curves depend primarily on the ratio of 
rotor conductor resistance to reactance. For instance, to 
obtain a greater locked-rotor torque, as well as a greater 
slip over the unable load range, rotor conductor resistance 
may be increased by decreasing the conductor cross- 
sectional area, or inductive reactance may be decreased by 
placing the bars closer to the rotor surface. On the other 



i cos e 



*-v 




rotor 



1 rotor 



B 



i cos e 




rotor 



rotor 



I cos e 



^r v,0, °' . 



i. 



s to tor 



A *rotor 

Figure 6.26.— Phasor diagrams of rotor and stator flux density for induction motor. 



140 



hand, an increase of conductor resistance will decrease 
overall motor torque and the stator current drawn during 
locked-rotor conditions. 

Single-cage rotors, as previously described, are the 
most rugged and the most used. Double-cage rotors use 
two conductors, one over the other, per rotor slot (fig. 
6.28A) and provide higher starting torques with higher 
load efficiency and lower running slip than the single 
cages (14). Here, the higher conductor would have high 
resistance and low reactance, while the lower set would 
have low resistance and high reactance. Double-bar rotor 
conductors are often susceptible to damage on loads with 
long accelerating times, lb overcome this problem, deep- 
bar rotors (fig. 6.28S) can be used. These have a thermal 
advantage in that the full conductor area is available for 
heat dissipation, but the design still approximates the 
performance of the double bar. Regardless of the design, 
the torque-speed curves are matched to the squirrel-cage 
rotor construction, which is fixed for a specific motor. 

In addition to rotor design changes, the actual values 
of breakdown and locked-rotor torque vary with the horse- 



300 



£ 250 - 



2 200 
DC 

o 

H 

Q 150 

< 

O 



i 



U. 





1 
--^^^ ^--Design D 


- 


^-Design C N. / \ 


. 


^" >, "**** ,, "-w x\ y^ \ \ - 




Design A — ^±.^ZS^\ \\ \ 








Design B ^^"^ ^\ \\| 


- 


^*-— " -""^^ Design F \\\\| 


- 


' 1 



100 - 



50 - 



50 100 

SYNCHRONOUS SPEED. % 

Figure 6.27.— Typical torque-speed characteristics for 
NEMA-design three-phase squirrel-cage motors. 



power, frequency, and speed ratings of the motor. Although 
the operating characteristics are a function of rotor imped- 
ance, the horsepower rating is mostly dependent upon the 
power (or kilovoltampere) capacity of the stator and rotor 
windings. As rotor losses are constrained to the rotor cage, 
rotor thermal capacity is limited. Therefore, motor designs 
that create large rotor currents, such as high-torque 
high-slip, may have intermittent time ratings or a limited 
number of allowed successive starts. Unless these con- 
straints are heeded, improper operation will burn out the 
rotor winding. 

The different rotor designs have led to a variety of 
speed-torque characteristics. To distinguish among the 
various types, NEMA uses a code letter system that 
signifies specific rotor constructions (8). Design B serves as 
the comparison basis for the motor performance of other 
designs and is often called the general-purpose motor. This 
design has relatively high efficiency even at light loads 
and a reasonably high power factor at full load. It has 
single rotor bars located rather deep in the core but with 
large-area slots for good heat dissipation. Starting cur- 
rents range from 4.5 to 5 times the rated full-load current. 
The design B motor has the broadest industrial applica- 
tion field. 

Design A has characteristics similar to those of design 
B, except that it has a higher breakdown torque. The rotor 
conductors are shallower, which decreases rotor reactance 
but increases the starting current, being five to seven 
times rated current. As a result, design B motors are often 
preferred over design A for large motor applications. As 
shown in figure 6.27, design A motors have the best speed 
regulation, as evidenced by the steep curve portion be- 
tween synchronous speed and breakdown torque (8). 

Design C motors have a double-cage rotor construc- 
tion that results in higher locked-rotor torque and lower 
breakdown torque than those of design B. Starting cur- 
rents are about 3.5 to 5 times rated current (8). These 
characteristics are well suited for conveyor belt drives and 
other applications that have sudden large load increases, 
but low or normal starting inertia. The motors are not 
suited for heavy high-inertia loads because the thermal 
dissipation is limited and high rotor current tends to 
concentrate in the upper bars (8). Accordingly, frequent 
starting of these motors can cause rotor overheating. 

Very high locked-rotor torque and high slip are found 
with design D characteristics. Design D's principal appli- 
cation is for high-inertia loads. The rotor is of high- 
resistance design with bars located close to the surface (8). 







bars 



uiah-reoctonce. 
ToVres-'Stonce 

bo< s 




/ 
// 

'/' 
'/' 
I. ■ 

II 

ill 



a 






1 TTT, /'I' 



Mi 



n. I 



\fe£' 



Rotor 
bor 



A Double-cage rotors B Deep-bar rotor 

Figure 6.28.— Other rotor-conductor designs. 



141 



Starting currents range from three to eight times rated 
load current. The motor is suited for heavy-duty starting, 
but again, the poor heat dissipation of the rotor design 
means that starting cannot be frequent. 

Design F has lower locked-rotor and breakdown 
torques than does design B. Design F motors also use a 
double-cage rotor with high resistance in both conductors, 
which reduces both starting and running current (8). The 
locked-rotor current is the lowest of all motor designs. 
Thus, design F motors are applied when starting-current 
limitations are severe and both starting and maximum 
torque requirements are low. The design, however, has 
poor speed regulation, low overload capacity, and usually 
low full-load efficiency. 

Induction-Motor Starting 

From the foregoing it can be seen that if an induction 
motor is started by directly connecting it to a power 
system, the momentary starting current can range from 
three to eight times the full load current. While this will 
not damage the motor, the high current can cause a 
significant disturbance on the power system, and, in some 
cases, activate overcurrent protection devices. However, 
most induction motors in mining applications are started 
by directly connecting them to the power system, espe- 
cially those within mining machines such as continuous 
miners. The system usually has enough impedance that 
protective devices can be set above the in-rush current to 
prevent nuisance tripping. This, however, is a major prob- 
lem, which is further discussed later in this chapter and in 
chapter 10. Full-voltage starting can usually be performed 
on 440- to 550-V motors up to 1,600 hp. NEMA standard 
magnetic starters for this range are shown in table 6.3 (3). 
The jogging service listed in the table refers to frequent 
stop-start or plugging (reversing under load) applications. 

As shown in figure 6.29, the across-the-line starter is 
simply three contacts driven by a solenoid, also called a 
contactor. Pressing the start button energizes the solenoid, 
which closes the M contacts. An auxiliary contact set (M b ) 
simultaneously closes and bypasses the start switch. 
Pressing the stop button deenergizes the solenoid. 



Above 1,600 hp (but sometimes lower), full-voltage 
starting becomes impractical even when the load con- 
nected to the motor can withstand the stress. Common 
methods for starting these large induction motors are 
shown in figure 6.30. In basic terminology, all these 
methods can be called reduced-voltage starting. In figure 
6.30A, an autotransformer is used to start the motor at 
reduced voltage (50% to 80% of rated), thus limiting 
starting current and torque. When almost at full speed, 
contactors quickly change the motor from the autotrans- 
former to the full-voltage supply. Primary resistor or 
reactor starting (fig. 6.30S) inserts fixed or variable im- 
pedances in series with the motor; these are shorted out 
after acceleration. For the wye-delta technique (fig. 6.30O, 
the motor is started as a wye connection, which places 
about 58% of the rated delta terminal voltage across the 
windings, limiting line current to 58% and torque to 35%. 
After acceleration, motor operation is with a delta connec- 
tion. Part-winding starting requires that the motor have 
two identical stator windings (fig. 6.30D). Starting uses 
only one winding and limits starting current to about 65% 
of normal, torque to 45%. After acceleration, the second 
winding is switched in. 

There are many systems that cannot take the shock of 
full-voltage starting. One instance is a conveyor belt drive 



Table 6.3.— NEMA class A standard starters for three-phase 
induction motors 





Continuous 


Maximum horsepower 


Maximum horsepower 


Controller 


current 


for normal service 


for jogging service 


size 


rating, A 


220 V 


440-550 V 


230 V 


460-575 V 


00 


9 


1.5 


2 


NAp 


NAp 





18 


3 


5 


1.5 


2 


1 


27 


7.5 


10 


3 


15 


2 


45 


15 


25 


10 


30 


3 


90 


30 


50 


20 


60 


4 


135 


50 


100 


30 


150 


5 


270 


100 


200 


75 


300 


6 


540 


200 


400 


150 


NAp 


7 


810 


300 


600 


NAp 


NAp 


8 


1,215 


450 


900 


NAp 


NAp 


9 


2,250 


800 


1,600 


NAp 


NAp 



NAp Not applicable. 



conductors 




3-phase 
motor 



-1 Start 
I Stop_|_ 

1 2^ 



O.L. 



L2 



-®-WM 



3-phase diagram Control circuit 

Figure 6.29.— Across-the-line magnetic starter. 



142 



3 -phase supply 
1-2 L 3 



l-l 

p 




A Autotransformer 



3- phase supply 
Li L 2 l 3 

o o o 

/ / / 



I *<\ 


/« 


1 
\ T 2 






< 











Tl Run 



Start 



# Primary reacter 



3 -phase supply 



3 -phase supply 




C Wye-delta 




Run 



D Part winding 



Figure 6.30.— Starting methods for induction motors. 



where the horsepower limit for full-voltage starting is 
perhaps as low as 50 hp. The wound-rotor motors described 
in the next section provide an alternative. 



WOUND-ROTOR INDUCTION MOTORS 

As mentioned earlier, the starting and running char- 
acteristics of an induction motor may be adjusted by 
varying the resistance-to-reactance (E/X) ratio of the rotor 
conductors. Instead of rotor bars and end rings, the wound- 
rotor motor has insulated windings much like the stator, 
with the same number of poles and windings placed in the 
rotor slots. The windings are usually connected in wye 
with the ends connected to three slip rings mounted on the 
rotor shaft. The brush and slip-ring circuit is completed 
through a wye-connected set of variable resistances, as 
shown in figure 6.31. Thus, the external resistance can be 
used to vary the speed-torque characteristics by changing 
the rotor R/X ratio. The stator of the motor is the same as 
for a squirrel-cage machine. 

A typical family of wound-rotor motor characteristics 
is illustrated in figure 6.32 (11). As external resistance is 
increased, the starting current is decreased and starting 
torque is increased. For a given shaft load, the reduction in 
rotor current will result in a speed decrease. Thus when 



starting a wound-rotor motor, a maximum resistance is 
inserted in the rotor circuit (R 9 curve). As the rotor 
accelerates, the resistance is reduced until the desired 
speed is obtained, or if full speed is required, the resis- 
tance is brought to zero (R x curve). Therefore the wound- 
rotor motor can be considered a variable-speed machine. 
Thermal considerations do place a lower speed limit on it, 
and for self-ventilated motors, continuous rated torque 
operation below 70% of rated full speed is not recom- 
mended (15). This lower limit may be reduced to 50% if the 
motor load is 40% of rated. 

Applications for wound-rotor motors include loads 
that require constant-torque, variable-speed drives or for 
which a sequence of slow-speed steps is needed to limit 
motor current during acceleration, such as for high-inertia 
or high-torque loads. Since they are suited to high-torque 
loads. Since they are suited to high-torque loads, these 
motors have found extensive used in the mining industry 
to operate crushers, grinders, ball and roller mills, con- 
veyor belt drives, and hoists. 

The automatic starting method for these motors uses 
definite-time acceleration where a series of fixed resistances 
are shorted out one at a time on a predetermined schedule 
(12). This step starter is shown in a simplified schematic in 
figure 6.33. When the starting sequence is initiated, all 
resistors are in series with the rotor winding; then the relay 



143 



3-phase winding on rotor 





External variable 
resistor 



Figure 6.31.— Schematic of wound-rotor induction motor 
showing external resistance controller. 



contacts 1A, 2A, and 3A are sequentially closed, resulting in 
four speed-torque characteristics. The last effectively shorts 
out the rotor winding. Since the sequence proceeds regard- 
less of motor speed, the method requires close coordination 
with motor characteristics {IS). The actual operation of the 
relays is discussed in chapter 9. 

Whether started automatically or manually, the 
wound-rotor motor continues to find application for the 
functions previously mentioned. However, for conveyor 
belt drives, these motors are now tending to be displaced 
by squirrel-cage induction motors equipped with solid- 
state starters (see chapter 14). Reasons for this change 
involve maintenance problems and a desire to eliminate 
the failures inherent with brushes, slip rings, and relay 
contacts. 



THREE-PHASE SYNCHRONOUS MOTORS 



250 




25 50 75 

SYNCHRONOUS SPEED, % 



100 



Figure 6.32.— Torque-speed characteristics for wound-rotor 
motor with stepped-resistance controller. 



The three-phase synchronous motor has a stator and 
rotor and is similar to the induction motor. The stator and 
stator winding have the same basic construction and 
purpose: to receive the power to drive a load (15). However, 
in this motor, the rotor consists of field poles connected in 
series, parallel, or series-parallel combinations and termi- 
nated at slip rings. The field windings are excited by an 
external dc source, the exciter. The number of field- 
winding poles equals the number of magnetic poles 
present in the stator. A sketch of a typical large synchro- 
nous motor is shown in figure 6.34 (12). 

Rotor field excitation is often supplied from a small dc 
generator mounted on the same rotor shaft, as dia- 
grammed in figure 6.35. Alternatively, dc supply can be 
obtained from a three-phase full-wave bridge rectifier, as 
illustrated in figure 6.36, or by a separate m-g set. 

Pure synchronous motors are not self-starting and are 
generally accelerated in the same manner as inductor 
motors. Salient-pole rotors commonly have a squirrel-cage 
winding (fig. 6.34) to produce the necessary induction 
motor action. Low-speed cylindrical rotors closely resem- 
ble a wound-rotor induction motor, but with five slip rings 



T3A T2AT1A 




<^f-Os.o-<rA>-« 
LB JCtC 



TS1A 

Atc 



TS2A 
CH 



"°$ 



-<gH 



H' T< 

3A 



Figure 6.33.— Simplified step starter using individually 
timed magnetic relays. 



Stator core 

Stator frame 

Stator core fastened to frame 
by dovetail keys on cross ribs 

Bolts, segment, and 
fingers for clamping 
stator core 

Dovetail 
core support 

// 
Stator coils 



Cage winding 
Cage bars 
Cage end ring 




Stator terminals 



Figure 6.34.— Sketch showing construction of salient-pole 
synchronous motor. 



144 



(15). As figure 6.37 illustrates, three rings are used for a 
wound-rotor circuit, the other two for the dc field (8). These 
cylindrical-rotor motors can provide high starting torque 
to accelerate high-inertia loads. The use of squirrel-cage 
windings, however, is intended only to accelerate the 



— «^ 

3- phase < ^ a 
supply l 





Synchronous 

motor 

armature 




"J_ Field 

/ /switch 



Field 
discharge 
resistor 



Figure 6.35.— Simplified diagram of synchronous motor us- 
ing generator for field excitation. 



motor rotor, not specifically to develop induction-motor 
torque to external loads. Some large synchronous motors 
are accelerated by a small induction motor mounted on the 
synchronous-motor shaft (12). The induction motor must 
have fewer stator poles than the synchronous motor in 
order to reach the required speed. 

Synchronous-Motor Starting 

Figure 6.38 demonstrates the general method of start- 
ing a synchronous motor (12). Pressing the start button 
energizes the CR relay, which in turn closes the CR 
contacts. One set of contacts electrically locks in the start 
sequence (which can be terminated by pressing the stop 
button), and the other set energizes the M relay. The M 
contacts close, and three-phase power is applied to the 
stator winding. This allows the machine to accelerate as 
an induction motor. In the simplest procedure for the 
motor shown in figure 6.34, the motor is allowed to 
accelerate to the maximum induction speed, where the 
slip between the stator rotating field and the rotor is very 
small. A switch is then closed manually to apply dc to the 
rotor field winding (figs. 6.35-6.37). A steady rotor mag- 
netic field is thus established that can lock in step with 
the rotating field of the stator. Thus, the rotor will turn at 
synchronous speed, which gives the motor its name. How- 
ever, if dc is applied before maximum induction speed is 



Li 
L 2 
L 3 



Slip ring 



Synchronous ^ otor rotQr 
motor stator 



sySSSus T P Brushes 



Mechanical 
load 







-/■vW\_ _i__i — II 



A-Y 

transformer 



y. :i 



Rheostat 
— vvv- 



1 T 



dc output 
voltage 



ac surge 
voltage 

suppression F ull-wave dc surge Smoothing 
rectifiers suppression filter 



Figure 6.36.— External solid-state supply used to provide field excitation. 



3 -phase damper 
rotor winding dc field 



Rotor i 

External rotor 
resistance 

-VWVy- 

~VAMr 

— vww- 



?Uf^Sv^--fc-- ,Bor 



-Slip 

rings 



Stator armature 

-o 




3- phase 
supply 



'o 'o 

dc supply 



Increase speed 

Movable shorting bar 
Figure 6.37.— Schematic of low-speed cylindrical-rotor synchronous motor. 



145 



achieved, the rotor may not pull into synchronization and 
severe vibration can occur, caused by repulsion every time 
a rotating pole passes a stator pole. As a result, most 
synchronous-motor starters do not rely on manual control 
but instead automatically excite the rotor field at the 
appropriate time. 

An approach widely used for automatic starting is 
synchronization based on frequency (12). This technique 
uses the voltage induced in the field winding during 
acceleration and before the dc is applied (1). Again refer- 
ring to figure 6.38, a resistor (R) and inductor (X) are 
placed across the field winding, with relay FR across the 
inductor. The inductance of the relay coil is selected to be 
much lower than that of X. Immediately after starting 
commences, a high-frequency potential is induced in the 
field winding, and the majority of current flows through 
the resistor and the relay coil because the inductance, X, 
exhibits high reactance. The FR relay opens the FR 
contacts faster than the interlock contact M b of relay M 
closes. As the motor accelerates, the frequency of the 
induced current decreases. When close to synchronous 
speed, the frequency has decreased to the point where 
most current flows through the inductor, and the voltage is 
reduced to the point where the FR relay cannot hold its 
contacts open. Consequently the FR contacts close and 
energize relay FS. The FS contacts then close to apply dc 
excitation to the field winding and remove the resistor 
from the circuit. 

Synchronous-Motor Torque 

Under load, the synchronous motor behaves much like 
a nonslip direct magnetic coupling. The rotor does not 
develop induction-motor torque; it is magnetically locked 
to the stator rotating field and is pulled around at basi- 
cally the same speed. The torque developed is dependent 
on the hold-in pole strength. Hence, the lock-in torque may 
be increased by simply increasing the dc supplied to the 
rotor field winding. If there is a load change, an instanta- 



3-phase supply 
LI L2 L3 




Stort 



Stop °T CR 



tCR 



R 



FR 



M 

-0-< 

5H 



neous speed change occurs but only for a few cycles, after 
which the rotor again attains synchronism (12). The use of 
a squirrel-cage rotor winding also helps to dampen out 
speed changes, and it is therefore often called a damper 
winding. 

A typical speed-torque characteristic for a synchro- 
nous motor containing a damper winding is shown in 
figure 6.39. Because the synchronous-motor portion can- 
not start itself, the starting torque comes from the damper 
winding. When the external loading does not exceed the 
pull-in torque value, the motor can be started and accel- 
erated to synchronous speed. However, if no significant 
external load exists and then a load equal to the pull-in 
torque is applied, the rotor will momentarily drop to about 
95% of full speed and then regain synchronism. During 
the loading and the momentary drop in speed, the rotor 
assumes a new position and continues to rotate at synchro- 
nous speed but a few degrees behind the no-load position (a 
in figure 6.40) (9). This sequence can occur for any applied 



NORMAL FULL-LOAD ARMATURE CURRENT, % 

ooooooooooo 
Ooujomomoiooioo 

OlflrrNNdn't^lfllBffl 



i — i — i — r 



i — i — i — i 



NORMAL FULL-LOAD TORQUE. % 



oooo OOC) OOn 

o S ? S o o N ^ » S 




10 
20 
30 








































1 












Torque< 


■\ 








/ 




40 














\ 


Current* 


-/ 




50 
60 
70 
80 
90 
















V 






/ 


















\ 




/ 














Storting torque' 


^ 


r 


















( 


y 




\ 
























) 






100 












i 




















Pul 


l-in 


r 

_ j 


— i 


i — 


^""^ 


Pu 


lOU 


1 



torque e..«»i..««« ^ torque 
Synchronous 

torque 

Figure 6.39.— Typical torque-speed characteristic for syn- 
chronous motor with damper winding. 



L11 L12 

D-C lines 




Figure 6.38.— Controller used to demonstrate general start- 
ing method for synchronous motor. 



Figure 6.40.— Effect of load on rotor position. 



146 



load up to the synchronous torque level, above which the 
restoration of synchronous speed is questionable. If the 
load requirements exceed the pullout torque, the motor 
loses synchronism, average torque drops to zero, and the 
motor stops (12). 

Generated Voltage 

After excitation has been applied to the field winding, 
the revolving magnetic field of the rotor cuts the stator 
conductors and induces a voltage in opposition to the applied 
voltage. Figure 6.41A shows an equivalent per-phase circuit 
of a synchronous motor that demonstrates the effect of the 
generated voltage on the electrical performance of the ma- 
chine. From Kirchhoff s voltage law (11), 



V T = V c 



+ m e + jix L = v c + v ? 



(6.11) 



where IR e = voltage drop due to effective resistance of 
armature (stator) winding, V, 
IX L = voltage drop due to inductive reactance of 

armature winding, V, 
Vj. = voltage supplied to motor, V, 
and V c = generated voltage produced by rotor field 
winding modified by armature reaction, V. 

Two important phenomena connected with synchronous 
machines (and some others as well) are immediately 
evident in the equation. Under dynamic loading condi- 
tions, if the load delivers a torque to the motor shaft, the 
rotor produces a generated voltage that is greater than 
that of the supply, and power is delivered back or regen- 
erated into the line. Secondly, if the supply voltage is then 
removed, the load acts as a prime mover, and V c will be 
generated as long as field excitation exists and until the 
load dissipates its energy. This last phenomenon is espe- 
cially important when the supply voltage is lost because of 
a short circuit, since the synchronous motor can deliver 
significant current to the malfunction (see chapter 10 for 
further information). 

Power Factor 

Another important effect results from the generated 
voltage. In_an ideal synchronous motor under no-load 
conditions, V c can be equal in magnitude and frequency to 
V t but 180° out of phase. Hence, with this ideal situation, 
the motor does not draw current (I). Obviously, practical 
motors have such losses as windage and friction, which 
cause a small shift in angular position, a, between the 
rotor jmd the rotating magnetic field. Here, the phasqrs V c 
and V T are no longer opposite in position, since V c is 
shifted clockwise by a as shown in figure 6.41.B. The 
change causes the motor to draw line current (I) to 
maintain the rotor in synchronism with the stator flux. 
Under heavier loading, a increases and the motor draws 
more current. Note that the rotor field actually does no 
work, and the dc energy supplied to maintain the field is 
dissipated as a small I 2 R heat loss. 

A change in the rotor field strength, say by adjusting 
the resistance _in figures 6.35, 6.36, or 6.37, changes the 
magnitude of _V C but _not its_angular position. The differ- 
ence between V T and V c , or Vxr m figure 6.41^ determines 
the angular position of motor current. When V c is adjusted 
to produce a motor current that lags applied voltage 
(fig. 6A1B), the motor is said to be underexcited. How- 



ever, increasing_the dc field strength with the same shaft 
load can shift Vxr such that the reactive component of 
current will change from a lagging phase angle to leading 
(fig. 6.41C). In this condition, the rotor field is termed 
overexcited, and the motor appears as a capacitive load. 
The leading power factor is one of the most outstanding 
features of a synchronous motor, as it can be used for 
power-factor correction. The ability to operate at unity 
power factor should be obvious; the field winding is 
referred to as normally excited in this case. 

Because the phase angle of operation depends upon 
both field excitation and motor load (angle a), the charac- 
teristics of synchronous motors are often represented 
graphically in a form called V-curves. Illustrated in figure 
6.42, they allow the selection of a field-excitation current 
for a load to produce a desired power factor (8). The lines 
drawn to show equal power factor are termed compound- 
ing curves (12). 



To dc 
exciter 




Figure 6.41.— Equivalent per-phase circuit of a synchronous 
motor (A) and phasor diagrams for (6) underexcited and (C) 
overexcited field winding. 



o 

Q. 



0.8 pf lagging 



Unity- 




-0.8pf leading 

Full load 
Half load 
No load 



Normal excitation 



►- Leading pf s 



Field current 
Figure 6.42.— V-curves for synchronous motor. 



147 



Applications 



Elementary Motor 



In the past, wide use was made of synchronous motors 
in the mining industry to take advantage of their constant 
speed and available leading power factors. Applications 
included ventilation fans, pumps, compressors, grinders, 
mills, and drive motors on m-g sets to provide power for 
dc equipment. However, static capacitors have now re- 
placed motors for power-factor improvement in almost all 
situations because of their flexibility and ease of installa- 
tion, and silicon rectifiers have supplanted m-g sets for 
power conversion. Nevertheless, one very important use of 
synchronous motors remains today: as the main drive 
motor in surface excavators. Here, one or more motors 
directly drive dc generators, which in turn power the dc 
motors serving the various functions on the machine. 
Figure 6.43 provides a plan view of a typical mining shovel 
where one synchronous motor drives three dc generators 
and the motor exciter (10). This subject will be continued 
at the end of the presentation on dc motors. 



Figure 6.44 illustrates an elementary two-pole dc 
motor with a one-coil armature. With the armature cur- 
rent flow as shown, the reaction of the armature magnetic 
field to that of the main field produces forces on conductors 
A and B, and the torque results in clockwise rotation. The 
commutator acts as a switch to reverse the armature 
current each time the conductors pass the neutral plane. 
To reverse armature rotation, the armature current flow is 
simply reversed. 

The two-pole motor is rather impractical. Torque is 
maximum when the plane of the armature conductor is 
parallel to the plane of the field, zero when at right angles. 
Figure 6.45 shows a four-pole armature with a four- 
segment commutator, but still with a two-pole main field. 
Here, motor torque does not drop to zero because an 
armature conductor is always under the magnetic influ- 
ence of the main field. Actual dc industrial motors have 
many commutator segments, armature conductors, and 



DIRECT CURRENT MOTORS 

The dc motor is the most versatile of all electrical 
machinery. On advantage over all the preceding motors is 
that its speed may be easily adjusted. The dc motors 
accounted for the majority of motors within the mine until 
the 1940's, when ac distribution systems started to replace 
dc. Induction machines then substituted dc motors, 
mainly because the ac-to-dc conversion of reasonably large 
power quantities was cumbersome. However, dc motors 
continued to hold prominence for some specific loads. In 
recent years, because ac-to-dc conversion is now very easy, 
dc motors have replaced some of their induction counter- 
parts. The reasons behind the extensive use of dc motors in 
mining will become apparent in the following paragraphs. 



•Field magnet 



Crowd joystick 



Hoist-and-swing joystick 



Boom foot lugs 



/ Swing \ f-""""j 

j machinery y j r -n Cj| 



Reeving winch 
(optional) 

Lubrication 
system 

(optional ) 



Hoist-propel 
transfer cabinet 



Right-hand 
operators' cab 
(optional ) 
J i 

Air 
compressor 




"*r— rp]Z 



Exci 
Crowd generator - 
800-hp ac motor ' 



Hoist-propel 
(^machine generator 



Figure 6.43.— Plan view of typical mining shovel showing 
m-g set. 




Mi|i|i|iW 

Figure 6.44.— Elementary two-pole dc motor. 




Brush 

Commutator 
Brush 



A Elementary motor 




B Torque output versus rotation 
Figure 6.45.— Elementary four-pole dc motor. 



148 



main field poles. The result is nearly constant torque 
output. 

As with the synchronous motor, the dc motor field 
does not do useful work. It merely provides the necessary 
medium for the armature windings to push against when 
developing rotary motion. In all but the very smallest 
machines, the field is supplied by dc through field wind- 
ings. The energy expended in these windings forms an I 2 R 
heat loss. 

Actual Motor Construction 

The essential motor parts are the armature (rotor), 
the commutator, and the main field frame and windings 
(stator). The armature is constructed of steel laminations 
pressed onto the shaft, with slots parallel to the shaft. The 
armature windings are placed in the slots and connected 
to the segments of the commutator, which is located at one 
shaft end. Carbon brushes, mounted on but insulated from 
the motor frame or one end bell, rise on the commutator 
segments. The main field windings surround laminated 
pole pieces that are bolted around the periphery of the 
motor frame. 

Interpoles (or commutating poles) are mounted be- 
tween the field poles (fig. 6.46), and the windings are 
connected in series with the armature (12). Their purpose 
is to improve commutation by opposing armature reaction, 
the distortion of the main magnetic field by the rotating 
armature field (fig. 6.47). The interpole windings produce 
a small magnetic field that opposes the main field in the 
same plane as the brushes. This reduces the magnetic 
field that is cut by the armature conductors undergoing 



Commutating field 



Commutating field 
winding 




Main field 
winding 



Compensating 

field 

winding 



Figure 6.46.— Cross-sectional sketch of dc motor showing 
interpole and compensating windings. 




Magnetic neutral (load) 
/ 

S } Magnetic 
'/neutral (load) 



Armature flux 



Resultant distortion of field flux 
produced by armature flux 



Figure 6.47.— Interaction between armature and main-field 
flux to produce main-field distortion. 



commutation (current reversal) and thus reduces brush 
sparking (3). 

The armature reaction can be further neutralized 
through the use of compensating or stablizing windings 
(12), which are placed in slots on the ends of the main field 
poles next to the armature (fig. 6.46) and are again 
connected in series with the armature. These windings are 
especially useful in motors intended for variable-speed or 
reversing operation. Without the compensation, armature 
reaction from large loads can neutralize the main field 
flux (12). 

Although interpole and compensating windings serve 
valuable functions in dc motor operations, the machines 
can work without them. As they can obscure the presen- 
tation of motor operation, these windings will not be 
included in the following discussion. 

Torque 

The torque developed by any electric motor is a 
measure of its ability to pull against a load. In dc motors, 
torque is a function of armature current and the magnetic 
flux density of the main field or (8) 



T = K#L 



(6.12) 



where T = motor torque, N-m (times 0.738 = lb-ft), 

$ = magnetic flux per main field pole linking 

the conductors, Wb, 
I a = total armature current, A, 

and K = a proportionality constant, N-m to WbA, 



and where 



K = 



L rated 



(I 



Aerated' 



A rated-"-* rated 



where T rated , I A rated , $ rated = rated value for torque, ar- 
mature current, and flux 
for motor, respectively. 

The above equation can be used to find the torque output 
from a machine, if the rated torque and the changes in 
armature current and the field flux are known. 

Motor Connections and Performance 

Exactly like dc generators, dc motors can be connected 
as separately excited, shunt, series, and compound. These 
connections are shown in figure 6.48. The performance of 
a separately excited motor is similar to that of the shunt, 
and its importance in mining applications is primarily 
with regard to motor control; thus, this motor will be 
discussed later. The speed-torque characteristics of shunt, 
series, and compound motors can change drastically de- 
pending on the connection. Typical curves are illustrated 
in figure 6.49 (8). 

Shunt Motors 

The shunt motor has the main field winding con- 
nected in parallel with the armature (fig. 6.48). Since the 
field winding is connected across the supply, its resistance 
must be rather high, but because of space constraints the 
armature windings have a much smaller resistance. When 
the motor is energized, armature current, I a , is limited 
only by its winding resistance and is thus much higher 



149 



I L =Ia 



W^f la 





Separately excited 



Il-VIq 



Shunt 



L=L=I» V^ 




Series Compound 

Figure 6.48.— Four connections for dc motors. 



100% speed 



Upper curves / 
are speed versus 
current 



Lower curves 
are torque versus 
current ^v 

100% torque 




ARMATURE CURRENT, % 



than field-winding current, If. However, as soon as the 
armature starts to rotate, its conductors cut the main field 
magnetic flux and a counterelectromotive force (cemf) is 
generated in the windings. This cemf opposes the applied 
armature voltage and begins to limit armature current. 
The opposition to current flow can be seen by applying 
Kirchhoff s voltage law: 



V T = 



V, + LR Q 



(6.13) 



Figure 6.49.— Typical characteristics for shunt, series, and 
compound motors of equal horsepower and speed ratings. 



where V T = supply voltage, V, 

V c = cemf induced in armature winding, V, 
and I a , R a = armature current, A, and resistance, B. 

The cemf is proportional to the speed of the armature, or 

V c <x n4>, (6.14) 

where n = armature speed, r/min, 

and $ = magnetic flux per main field pole, Wb. 

As the armature accelerates, the cemf rises, and the 
armature current drops. Yet, according to equation 6.12, 
motor torque decreases. A final speed is reached when the 
cemf is almost equal in magnitude to the supply voltage. 

If the motor is unloaded, the difference between the 
terminal voltage and the cemf will allow only enough 
armature current to overcome friction, winding, and core 
losses. Under motor loading, the armature slows down, 
cemf decreases, and more current enters the armature. 
However as shown in figure 6.49, the speed of the shunt 
motor remains relatively constant from no-load conditions 
up to 100% rated and slightly beyond. The speed can be 
easily adjusted by changing a resistance in series with the 
field winding. From equation 6.14, weakening the field 
flux by decreasing field current increases motor speed. Yet, 
for a constant field flux, torque varies linearly with 
armature current (that is, T oc I a ). 

If across-the-line starting was attempted with the 
shunt motor shown in figure 6.48, the cemf would proba- 
bly not build up fast enough to limit armature current to 
a safe value, and hence, damage to the commutator, 
brushes, and the armature winding could result. For this 
reason, a starting resistance is used in series with the 
armature (fig. 6.50A) for all dc motors except those of 
fractional horsepowers. The resistance is usually selected 
to limit armature current from 150% to 250% of rated 
current depending on the starting torque required. The 
shunt winding is always connected across full line voltage 
when starting so less armature current is needed to 
develop the rated torque. 



Variable starting resistor 




Speed- 
adjusting 
rheostat 

Shunt 
field 



Variable starting resistor 




Variable starting resistor 

+ 







A Shunt motor 



B Series motor 



C Compound motor 



Figure 6.50.— Simplified dc motor schematics with starting resistances. 



150 



In mining, manual controllers are found on many dc 
machines. These are available in three general forms: 
faceplate, multiple-switch, and drum controllers. Sche- 
matics for these are shown in figures 6.51, 6.52, and 6.53. 




-Holding 
coil 



Figure 6.51.— Faceplate manual starter. 



Circuit breaker 



Overload trip 




±l_ine 

Disconnect with 
cover interlock 



Shunt fields 



Both switches 
must be in 
off position 
to open cover 



w Series field 



Figure 6.52.— Multiple-switch starting. 



I 2 3 4 




Shunt field 
Figure 6.53.— Drum-type starter. 



The faceplate starter is often used with small station- 
ary dc motors. The level is advanced (to the right) in steps, 
momentarily stopping at each position to allow the motor 
to accelerate, until the resistance is removed. A holding 
coil then maintains the lever in the last position. A spring 
is used to return the lever to the off position during a 
power failure or if the lever is left in an intermediate 
position. 

One method of multiple-switch starting, shown in 
figure 6.52, uses two double-pole, single-throw switches. 
The upper switch is closed first, energizing the shunt field 
through a 100-B resistor. This allows the main field flux to 
build up to some extent before the armature is connected 
to the line. Initial inrush current is thus reduced, which 
helps to prevent brush arcing and the possibility of com- 
mutator flashover. The lower switch is then closed, ener- 
gizing the armature through a second resistor, and the 
motor accelerates. The armature resistor remains con- 
nected during running. The two line switches are mechan- 
ically interlocked so the upper switch must always be 
closed first. A variation of this technique is to use relay 
contacts or contactors to supply main field excitation, 
insert the starting resistance, then bypass the resistance. 

Drum controllers (fig. 6.53) are frequently used on 
mine locomotives but are also found on some dc mining 
machines. A handle-controlled rotary shaft is connected to 
the switch segments indicated by dark lines in the figure. 
These segments are of various lengths so contact with the 
stationary contacts can be made at different intervals. 
When starting, the Ml and M2 contacts engage first, 
energizing the shunt field and inserting all resistors in 
series with the armature. The resistors are then removed 
one at a time by advancing the controller. Although not 
shown, an additional drum or reversing controller is 
usually available to reverse armature current and thus 
motor direction. 

The use of fixed resistance starting has widespread 
application in mining. Here the starting resistance re- 
mains in series with the armature for running. An in- 
stance would be a small dc motor, such as a pump in a 
remote location. The resistance gives poor speed regula- 
tion, but the motor can be started unattended. 

Dynamic Braking 

If a shunt motor is running under load and the 
armature circuit is opened, the inertia of that load will 
drive the machine as a dc generator. Dynamic braking 
simply connects a resistance across the armature to dissi- 
pate the available energy and decelerate the load (fig. 
6.54). The braking action is most effective at high arma- 
ture speeds, becoming negligible at low speeds. The value 
of resistance, R, is selected from (12) 



R _ V c - I a R a 



(6.15) 



where V c - I a R a = armature voltage at start of braking, 

V, 
and I = dynamic braking current, depending 

upon desired braking level, A. 

The normal value for I is 150% of rated motor current but 
I may be as high as 300% for quick stopping. 



151 



Series Motor 

The armature and main field winding are connected 
in series and both carry load current in a series motor. The 
magnetic flux, $, now produced in the main field winding, 
is proportional to the armature current. Thus, motor 
torque varies as the square of armature current (T <x I 2 a ). 
Furthermore, the main field strength will change with 
load, causing a speed decrease with increased load. 

When a series motor is started, cemf builds up as the 
armature speed increases. During the initial acceleration, 
the cemf is small, armature and field current are high, and 
the torque is very high. When the curves in figure 6.49 are 
compared with the material presented earlier, it can be 
seen that the starting torque of the series-wound motor is 
higher than that of any other motor type. Because of this, 
it is often said that the series machine has the best 
traction or starting characteristics. Thus, it is the most 
used motor for traction purposes in mining; examples 
include locomotives, shuttle cars, and diesel-electric 
trucks. A problem with this motor, however, is that at light 
loads, motor speeds may become excessively high; there- 
fore, series motors must be directly connected to loads that 
cannot be removed freely. Otherwise, the motor may race 
to destruction. 

The method for starting the series motor is similar to 
that for shunt machines. The arrangement of the starting 
resistance is shown in figure 6.50B. A direct application of 
contactor-controlled multiswitch starting of a traction 
motor in a mining machine is illustrated in figure 6.55 (7). 
After the power-source contacts (Ml) close, the motor is 
accelerated with both resistances in series with the arma- 
ture. The same control that activated the Ml contacts 
simultaneously energizes a definite-time relay. After a 
preset time (about 1.0 s), the relay closes its contacts, and 
that in turn energizes the M6 contactor, which shunts its 
starting resistor. The M7 contacts can be used to provide a 



L1 



L2 



Shunt field 



Run 



Armature 



(i 0/ ArmaTure 

^—O — i— ■ 



Brake 



-CD- 



Resistance 



Figure 6.54.— Simplified diagram of dynamic braking ap- 
plied to shunt motor. 



second step before full speed is obtained or can be used to 
enable two-speed operation of the motor. In the latter case, 
control circuitry is arranged so that the M7 contacts 
cannot close before the M6 contacts. Figure 6.56 shows a 
one-step starting arrangement with the addition of 
forward-reverse control (contacts IF and 2F close for for- 
ward, 1R and 2R for reverse) (7). 

The procedure for dynamic braking is identical to that 
already described, with the exception of excitation for the 
main field. The simplified circuit in figure 6.57 is one 
approach and shows the switches closed for motoring (12). 
Upon dynamic braking, the switches place the armature, 
series field, and braking resistance in a loop circuit. The 
series-field connections are reversed to maintain current 
flow in the same direction. 

Compound Motors 

Compound motors have both shunt and series field 
windings installed on the same poles. The series winding 
may be differentially or cumulatively compounded, that is, 
subtracting from or adding to the magnetizing force of the 
shunt field. This causes either reduced or increased arma- 
ture speed with load. Only the cumulative compound 
motor characteristics are shown in figure 6.49. 

Cumulative compounding gives greater torque than is 
possible with the simple shunt motor, because of the 
greater amount of main field flux available (8). The 
increased flux, however, causes the speed to drop off more 
rapidly than for a shunt motor, but not as much as for a 
series motor. Therefore, the cumulative compound motor 
will develop a high torque with any sudden increase in 
load, but at light loads it will not run away because the 
shunt field provides a constant field flux. These motors are 
often applied to loads requiring high starting torque but 
fairly constant operating speed under normal conditions. 
Thus, cumulative compounding combines the characteris- 
tics of both series and shunt motors. 

The differential compound motor produces torque that 
is always lower than that of the shunt motor (8). The 
amount of series winding can be adjusted to offset any drop 
in speed as loading increases, or it may be sufficient to 
give a slightly higher speed than normal at full load. A 
motor having constant speed from full load to no load is 
called flat compounded, while that with slightly higher 
speed than normal is called over compounded. 

Again, armature current is traditionally limited by 
resistance when starting. Figure 6.50C shows the process 
in elementary form, and figure 6.58A illustrates an actual 
application for a mining n.ac 1 ne hydraulic pump motor 
(7). It can be seen that the wo circuits are almost identi- 
cal. A fixed starting resistor is used for acceleration and as 
in figure 6.55, the resistor is shunted by the M3 contacts 
after a definite til period (usually 1.0 s). A semi- 
automatic variation of this scheme is illustrated in figure 



M 
CO 



Storting 
■ M 1 resistor 

jnnector j | | I 



M, 



Commutator 
field 



Starting 
resistor 




Overload 
coil M 1 

■Ar— IK 



Armature fj e id 

6 ""7 

Figure 6.55.— Two-step resistance starling of series-wound motor. 



M- 



152 



6.58B: semi-automatic means that under certain condi- 
tions the starter requires some attention. The accelerating 
contacts (A) are as before, but the contactor coil is placed 
across the armature circuit. As cemf increases during 
acceleration, the voltage across the coil causes contact 
closure at the proper time. 



+ O 



Series 
field 



Commutator 
field 



Starting 
resistor 




Dynamic braking employs a resistance to dissipate 
energy generated in the armature, involving either the 
series field (fig. 6.57) or simply the armature itself (fig. 
6.54). 

Ward-Leonard System 

For large-motor applications, the Ward-Leonard sys- 
tem provides one of the finest techniques for controlling 
motor speed over a wide range and in both rotational 
directions (3). Two specific examples where it is used are 
mine hoists and surface excavators (2, 4). Figure 6.59 
illustrates the basic system. The dc generator is driven at 
constant speed, typically by a synchronous motor, but 
some systems employ induction-motor drives for smaller 
horsepower applications. The generator and motor field 
windings are separately excited (see exciter in figure 6.43), 
and the motor is excited with a constant field current. 

Because the main field of the motor is constant, the 
speed is directly proportional to its armature cemf (V m ). 
The magnitude of V m is directly dependent upon the 
generator output voltage (V g ) less i a R a , where i a is the 



L1 



Run 



^-o 



Brake 




-O — • 



Resistance 



Figure 6.56.— Forward-reverse switching of series-wound 
motor. 



Figure 6.57.— Dynamic braking applied to series-wound 
motor. 



Starting 
resistor 




Main line 



Commutator Series 
field field 



1 • Armatur 




-TWA. 



Overload 
coil M 

AH 



Shunt field 



Starting 
resistance 



A 
Ht- 



Shunt 
field 



Series 
field 



^ Starring 

o reloy 



Figure 6.58.— One-step starting of compound-wound motor. 



153 



Generator 



i — VW 



— ^ Motor 
-W\ 1 




ojg = constant 



+ Vfg - 

Figure 6.59.— Basic Ward-Leonard system. 



armature current and R a is the combined resistances of 
the generator and motor armatures. As a result, excellent 
control of all motor speeds and both acceleration and 
deceleration is obtained by adjusting the generator field 
strength. The generator field-winding resistance is high, 
and so the required level of control power is relatively low. 
Motor reversing is obtained by changing the current 
direction in the generator field. Braking is performed by 
reversing or reducing the generator field current. 



MINE MOTORS 

Many mining uses for industrial motors have been 
covered so far in the chapter; this section serves to clarify 
some additional applications, but mostly for underground 
mining equipment. 

Applications 

Mine motor functions can usually be divided into two 
groups: auxiliary and face (17). Auxiliary motors are 
employed for fans, pumps, conveyors, hoists, compressors, 
and other vital functions in mines aside from the actual 
process of mineral extraction. These operations commonly 
call for direct use of general-purpose industrial motors, 
and as their loads are often well defined and continuous, 
the motor characteristics covered so far are applicable. 
Face motors are associated with mining equipment, such 
as continuous miners, shuttle cars, loaders, roof bolters, 
and locomotives, where they are mounted in the machine. 
Their duty usually involves cyclic or random loading as 
well as the possibility of shock loading. The result is 
higher electrical and mechanical demands than those 
placed on equipment in other industrial applications. 

The horsepower rating for a motor is based on the 
maximum winding temperature for continuous duty or 
intermittent duty. The temperature rise parameters have 
already been covered, but the meaning of a duty cycle has 
not. Continuous duty is quite obvious and refers to a 
substantially constant load (torque) over an indefinitely 
long period. Intermittent duty, however, means that load- 
ing is at alternate intervals of load and no load (motor 
running idle); load and off; or load, no load, then off (9). 
Each portion of the cycle is equal and the time interval is 
specified. In some cases, face motor intermittent duty is 
given a definite time interval of 15, 30, or 60 min, but it is 
often just listed as "mine duty" (6, 17). Tsivitse (17) states 
that a very successful horsepower rating for face motors 
has used both the continuous rating and the 60-min 
rating. The continuous duty is matched to the average or 



rms requirements of the load, and peak horsepower load- 
ing is limited to the 60-min value. The rms value for 
horsepower can be defined as 



hp r 



EAt ; 



(6.16) 



where hp 4 = mean horsepower during time segment Atj, 
and EAtj = total time interval. 

The ac motors in mining machines are normally four 
or six pole with synchronous speeds of 1,800 and 1,200 
r/min, whereas dc motors often have comparable base 
speeds of 1,750 and 1,175 r/min (6, 17). These speeds are 
high enough to provide adequate horsepower but low 
enough to have reasonable reliability. Series-wound mo- 
tors for traction are built to withstand rotation up to 6,000 
rpm, such as might occur during maintenance. 

Table 6.4 contains a listing of common applications for 
different motor designs to accommodate the various func- 
tions found in mining equipment (5, 17). Some additional 
information is warranted. The locked torque of traction 
motors is set so that the wheel or crawler-tractor treads 
will lose traction before the motor stalls. Ac motors that 
are mechanically paralleled, as for coal cutting with a 
continuous miner, are often sequence started with a delay 
to limit starting currents. Further, the high-slip charac- 
teristics mentioned in the table are for load sharing as 
well as to limit the rate of torque rise during shock 
loading. The dc motors used in load sharing often have 
matched speed-torque characteristics. Otherwise, the mo- 
tors are compounded with a differential field that is in 
series with the armature of the second motor. 

Table 6.4.— Common motors for mining equipment 



Function 



dc 



Traction (direct coupled) . 
Hydraulic pump 



Design D Series-wound. 

Design B or A Compound or 

stabilized series. 2 



Conveyor 

Loading arm. 

Cutting 



Design C ... 
( 3 ) 

f) 



Compound. 
Heavily 

compounded. 4 
Do. 4 



1 NEMA designs tor induction machines. 

2 Speed regulation from no load to full load from 10% to 15%. 

3 Similar to NEMA design A but sometimes with higher locked-rotor and 
breakdown torque and higher slip per torque value. 

4 Speed regulation from no load to full load from 30% to 35%. 



Actual Equipment Operation 

Because mining equipment operates at the tail end of 
the distribution system, voltage drop becomes an impor- 
tant factor in the selection and utilization of motors. This 
is more critical for ac equipment than it is for dc, and this 
section will discuss some ramifications for two types of 
machines. 

Continuous Miners 

Of all electrical equipment used in U.S. underground 
coal mines, the continuous miner is the most concentrated 
simple load. This machine is the heart of present under- 
ground coal mining systems from both a production and 
electrical standpoint; hence, determining the load de- 
mands it makes on the system, or the load factor, is of 



154 



great importance. The machine load factor can be defined 
as the ratio of actual power consumption to rated motor 
power. The rated power for the squirrel-cage induction 
motors used on ac continuous miners is set by the manu- 
facturer for one motor or a combination of motors. The 
motors may or may not be built to NEMA standards. 
Regardless, torque and power are the only common ratings 
available to judge motor utilization. Horsepower is di- 
rectly proportional to the product of motor speed and 
torque, and this power rating can be employed to deter- 
mine three-phase motor performance. The load factor can 
be used not only to investigate the effective operation of a 
particular machine, but also to compare different equip- 
ment of a specific type, no matter what the rated power. 

At low load conditions, motor efficiency and the power 
factor drop off rapidly. Since the motor functions on the 
steep portion of the power-factor curve, a small load 
variation will cause a relatively large motor current 
fluctuation. This can produce detrimental current peaks 
and stresses in trailing and feeder cables, particularly 
where conductors have marginal size. Poor power-factor 
operation requires correction capacitors or results in util- 
ity company penalties. 

To analyze the power factor of continuous miners, a 
recent study (13) investigated the actual operation of 26 
different ac continuous miners. These machines had utili- 
zation voltages of 440, 550, and 950 V, and total rated 
motor powers from 100 to 535 hp. All were operating in the 
Appalachian coalfield, with production ranging from 50 to 
770 raw tons per shift. Average load factors were deter- 
mined for each machine and particular attention was paid 
to the cutting-and-loading machine cycle because here 
power consumption is the most demanding. The average 
cutting load factors ranged from 0.26 to 1.17 and averaged 
0.52 for the measured machines. It is significant that this 
average load factor is much less than the assumed design 
level of 0.85 that has been popularly used in the industry. 
When employing all hydraulic, mechanical, and electrical 
machine components, a 0.60 load factor might be consid- 
ered satisfactory for the continuous miners studied. 
Hence, the implication was that many were being used 
inefficiently. However, drawing conclusions about machine 
efficiency and utilization based only on the load factor 
could be misleading because of the numerous factors 
involved. 

Many of the low to moderately powered machines in 
the study had higher-than-average load factors, and some 
were considered to be totally adequate. During field mea- 
surements, close attention was paid to the performance of 
the machine operator, and in almost every case it was 
found that the operator was pushing his machine as much 
as possible during sumping (the cutting cycle), because 
traction approached full slip. From the standpoint of 
adequate employment, some continuous miners could be 
termed overpowered, particularly in the case of high- 
powered machines cutting friable coal. 

When a high-horsepower machine (500 hp or more) is 
used on a low- voltage system, the demand for large current 
can create considerable trailing-cable voltage drops and a 
machine voltage condition that not only hampers opera- 
tion but reduces the safety levels of the system on which it 
operates, perhaps by causing a poor power factor or sub- 
stantially reducing motor torque capabilities. In the study, 
good machine voltage conditions (that is, as close to ± 10% 
as practical) almost invariably resulted in good motor load 
factors. To obtain the favorable operation required not only 



good voltage but a strong utilization system, that is, using 
the largest practical trailing-cable conductors and shortest 
practical trailing-cable lengths. When distribution voltage 
regulation was bad, poor machine load factors also oc- 
curred. The situation was most evident on 4,160-V distri- 
bution systems that had been extended beyond their 
limits. More information on these subjects is presented in 
chapters 8, 12, and 13. 

Traction Locomotives 

A specific case study that was associated with the 
preceding work involved measurements on main-line trac- 
tion locomotives. The results of the study can be applied to 
all series-wound motor traction. The company involved 
was experiencing numerous motor armature failures on 
their locomotives— up to 47 in 1 yr. Two avenues were 
explored to determine the problem: inspection of the failed 
motors and electrical measurements on a typical operat- 
ing machine. Examination of the motors showed that the 
commutators were heavily pitted and charred, which was 
a direct indication of overloading. Subsequent electrical 
measurements substantiated this suspicion. The locomo- 
tive apparently experienced severe continuous stress that 
caused abnormally low motor voltage every time it en- 
countered a particular curve located on a steep upgrade. 
When provided with this information, the mining com- 
pany was able to reduce the locomotive trailing load, and 
the motor failures diminished. 

In this case, the very low motor voltage provided the 
clue to identifying the problem. Unlike ac motors, dc 
series-wound motors can still operate under low voltage, 
although their control circuitry might not function prop- 
erly. Here, the low voltage indicated high current because 
the trolley system was well maintained and had adequate 
capacity and properly spaced rectifiers. 

There is another method that indicates if the motors 
within a vehicle are being overstressed while performing a 
specific duty cycle. Every manufacturer supplies charac- 
teristic curves for its mining equipment. The example in 
figure 6.60 shows motor characteristics for a small loco- 
motive using two series-wound motors. The technique 
consists of finding the current needed by the job and using 
the characteristic curves to compare the needed value with 
the maximum value allowed per motor. The classical 
method employs rms current and makes the following 
assumptions: 

• The vehicle operates under constant velocity while 
performing a specific function; 

• The motors heat during acceleration and cool for 
deceleration; and 

• The ampere rating for the motors stabilizes after 8 
h of operation. 

The method requires complete knowledge of the entire 
duty cycle, which in mining is the repetitive process that 
places individual demands on the machine. For instance, 
the track profile for a locomotive under loaded and then 
unloaded conditions can be divided into segments of equal 
demand, such as the grade for a specific haulage distance. 
For each portion of that duty, the torque or tractive effort 
demand must then be found. Stefanko (16) contains meth- 
ods to calculate this input information. 

The characteristic curves are then used to find the 
current demand and the actual time the machine operated 



155 



at that current for each portion of the duty cycle. For 
instance, consider that the locomotive of figure 6.60 is 
operating on 600 ft of track with a + 0.5% grade and has a 
tractive effort per motor of 1,020 lb. From the tractive 
effort curve, the current demand is 81 A, and using the 
speed curve, the machine speed is about 8.1 mi/h. By 
simple calculations it can be found that the locomotive 
would take 0.84 min drawing 81 A to move the 600 ft. The 
technique is continued until all times and currents are 
known for each duty cycle portion. 

A problem occurs in the above procedure when the 
speed obtained from the curves is greater than that 
allowed. For example, assume that the next portion of the 
track profile has a length of 2,400 ft at a - 0.5% grade, and 
the tractive effort for each motor has been found to be 340 
lb. Referring to figure 6.60, current is 42 A, and speed is 
12.5 mi/h. However, say that the maximum allowable 
speed is 10 mi/h. In this case, the locomotive would be 
commonly operated on-off, on-off, and so on, to maintain 
but not exceed 10 mi/h throughout the haulage portion. 
The time the motors are on and off can be calculated 
precisely by 

1. Finding the time required to travel the distance at 
maximum allowed speed (2,400 ft at 10 mi/h yields 2.73 
min), 

2. Determining the time it would take at the speed 
found from the curve (2,400 ft at 12.5 mi/h gives 2.18 min), 
and 

3. Subtracting the results of item 2 from 1 (0.55 min). 

Item 2 provides the time the motors are on (42 A at 2.18 
min), while item 3 gives the off time (zero current for 0.55 
min). 

With all currents and times known, including those 
times at zero current, the following equation provides the 
rms current demanded by the duty cycle: 



9,000 




18 ■ 
17 - 


8,000 




16 - 

15 - 


7,000 




14 - 
13- 


6,000 




12- 




-C 


II - 


5,000 


- E 


10 - 




o 

Ld 


9- 


4,000 


</} 


8 - 
7 - 


3,000 




6 - 
5 - 


2,000 




4 - 
3 - 


1,000 




2 - 







- 




J I I ' I 



20 40 60 80 100 120 140 ISO 180 200 220 240 260 280 
MOTOR CURRENT, A 

Figure 6.60.— Typical characteristic curves for each motor 
in traction locomotive (8-ton, 250-V motor, characteristic 
curves on 250 V; pinion, 13 teeth; gear, 69 teeth; wheel 
diameter, 29 in). 



the 25° C ambient. The base temperature closely fits the 
typical conditions found in underground operations. 

As mentioned earlier in the chapter, the allowable 
temperature rise is effective to elevations of 3,300 ft (1,000 
m); above this, the allowed is reduced 1% for every 330 ft 
(100 m), or 



t _ /j*jj JaW 

Arms y J,. 



(6.17) 



% reduction = 



elev - 3,300 
330 



(6.18) 



where 1^ = current demand for each duty cycle portion, A, 
tj = time involved for respective current demand, 
min, 
and Ij^s = effective current demand for duty cycle, A. 

For additional functions performed by haulage locomo- 
tives, the following factors can be assumed: 

1. Switching operations have zero current demand but 
one-quarter of the actual time is applied in the above 
summation. 

2. If the locomotive is used to load and unload its cars, 
the maximum tractive effort is employed for the loading 
process, the minimum for unloading. One-half the actual 
time involved for each is used for the effective time (t { ). 

3. Delays are taken as zero current and zero time. 
Normal delays are assumed to not allow effective motor 
cooling. 

Mine motors are presently standardized at a 90° C 
allowable temperature rise based on a 25° C ambient 
temperature (17). Older motors, as in figure 6.60, may 
have a 75° C temperature rise limit but are still based on 



In addition, for maximum ambient temperatures exceed- 
ing 25° C, the allowable rise must also be reduced by the 
difference above the base temperature. For example, a 
motor with 75° C temperature rise insulation, operating 
at 6,600 ft elevation in 30° C ambient temperature, has 
only a 62.5°C allowable temperature rise. 

Consequently, if the locomotive is operating at 3,300 
ft or less in a maximum ambient of 25° C or less, the 
characteristic curves can be used directly to find if the 
duty cycle demands exceed that allowed. In other words, 
the rms current found by equation 6.17 can be compared 
with the time-to-rise temperature curve. If the resulting 
time is greater than the actual time involved, the locomo- 
tive will work under that duty cycle. For example, if 1^,,, is 
80 A, 7.5 h of operation is allowed (fig. 6.60). 

However, if the allowable temperature rise must be 
reduced, the manufacturer curves can no longer be used 
directly and must be corrected. Fortunately, the time-rise 
curve is very nearly parabolic. Thus, any allowable tem- 
perature rise curve can be closely approximated by a 
straight line through two points plotted on log-log paper, 
with the two axes representing motor current and opera- 
tion hours. This process can be time consuming. Using the 
parabolic relationship, the following formula also gives 



156 



the allowable effective (rms) current for the total operating 
time when curve correction is necessary: 



Rotor 
winding 



. m [ln(Hr)- Waind^ -ln(I 2 )] . _ . 
ln(l) = - + ln(l 2 ), 



(6.19) 



where W = In 



Z = In 



[(I 2 2 H 2 ) 1/2 (T D /T C )] : 
T 2 

q x 2 H,) 1/2 (T D /T C )]' 
Ix 2 



H T = total operation time for motor, h, 
T c = rated allowable temperature rise, °C, 
T p = corrected allowable temperature rise (due to 
elevation or ambient temperature), °C, 
H 1 ,l 1 = a point taken from manufacturer's time-to- 
rise temperature curve, h, A, 
H 2 ,I 2 = a second point taken from curve, h, A, 
and I = allowed rms current for total operation time 
H T , A. 

If the allowable current from equation 6.19 is less than 
Ij^g from equation 6.17, the motor is overstressed for that 
duty cycle. Even though the foregoing was applied to dc 
motors in traction locomotive, the same concepts can be 
adapted to any mine motor application. 



SINGLE-PHASE MOTORS 

Although the vast majority of mine motors are three- 
phase and dc, single-phase motors do find widespread use 
for auxiliary functions aside from the mining process. As a 
general rule, single-phase induction motors have one run- 
ning speed and require a separate means for starting 
rotation, usually a separate stator or starting winding. 
Motors are classified by their starting method. The most 
used techniques are split phase and capacitor start, which 
will be discussed briefly in this section. 

Rotating Stator Field 

When a single-phase ac voltage is applied to one 
stator winding, the current flow produces a magnetic field 
with a resultant direction that alternates on a line, as in 
line OP in figure 6.61. If a squirrel-cage rotor winding is in 
the stator field, a voltage will be induced in the rotor 
conductors, but the current produced will create a mag- 
netic field that coincides with the stator field (fig. 6.62). As 
no magnetic interaction occurs, no torque is developed, 
and the rotor remains stationary (9). 

If the rotor is moved by some means, the rotor conduc- 
tors cut the stator magnetic field, and the induced voltages 
are in phase with the current through the stator winding. 
However, the rotor winding impedance appears as almost 
pure inductance, and rotor current will lag the induced 
voltage by almost 90° (fig. 6.63). Thus the rotor magnetic 
field is now 90° from the stator field and is termed a 
cross-magnetizing field (9). The rotor and stator fields 
combine to produce a resultant field that rotates at syn- 
chronous speed. The cross field strength is proportional to 
the rotor speed, and is about equal to the stator field 
strength at synchronous speed. The same operational 
principles that have been given for three-phase induction 




Stator 
winding 



Figure 6.61.— Stator field of two-pole, single-phase induc- 
tion motor. 



Rotation 




Figure 6.62.— Rotor field of stationary two-poie, single- 
phase induction motor. 



■ Generated rotor voltage 
Stator current and flux 
Rotor current and flux 




Figure 6.63.— Phase relationships between stator and turn- 
ing rotor. 



motors also hold for single phase; slip must always exist 
between the rotating field and the rotor. Because of the 
cross field, the slower rotor speed causes the rotating field 
to pulsate. Accordingly, vibration and noise are inherent 
with single-phase induction motors. 



157 



Split-Phase Starting 

Split-phase motors have two stator windings con- 
nected in parallel, as shown for the two-pole motor in 
figure 6.64. The impedance of each winding is such that 
the currents through them are out of phase. One winding, 
the auxiliary or starting winding, is usually constructed of 
small-gauge wire and has high resistance and low reac- 
tance. The running or main winding has a heavier gauge 
conductor so the winding is of low resistance and high 
inductance. When energized, the phase angle between the 
currents through the two windings is only about 30°, but 
this is enough to produce a rotating magnetic field. The 
rotating field pulsates, and starting torque is small. 

Once the motor is started, the rotor cross field is 
produced. Thus the starting winding is no longer needed, 
and it is usually disconnected when the rotor speed 
reaches 70% to 80% of synchronous (9). A centrifugal 
switch mounted on the rotor shaft is almost always used 
(fig. 6.65). 

The starting direction determines the final rotating 
direction. Unlike three-phase motors, single-phase induc- 
tion motors must be stopped and the starting-winding 
connections reversed, then reenergized to produce a rotat- 
ing field in the opposite direction. 

Capacitor-Start Motors 

The capacitor-start motor also has two stator wind- 
ings. The main winding is arranged for direct connection 
to the power source, and the auxiliary winding is con- 
nected in series with a capacitor. With this arrangement, 
the currents through the two windings can be as high as 
90° out of phase. Hence, starting torque can approach 
100% of rated (9). Typically, the starting winding is 
disconnected at 70% to 80% of synchronous speed. A 
centrifugal switch or a relay sensing current through the 
main stator winding may be used (fig. 6.66). Apart from 
the high starting torque, the operation of capacitor-start 
motors is basically the same as split phase. However, 
popular split-phase motors have an upper power limit of 
1/3 hp, whereas capacitor-start machines can be obtained 
up to 10 hp. 

This chapter has introduced the operation and char- 
acteristics of the motors in common use in the mining 
industry. Although elementary in nature, the contents of 
the chapter should not be discounted. The electrical power 
systems in and about mines have the purpose of ade- 
quately serving motors. If the characteristics of these 
loads are not precisely known, it is doubtful that a safe 
and effective mine power system can be achieved. 











i 




> Starting winding 

> >4 and poles 




» 


N 






r^ f^ ^ 


Rotor 










•it 




s 






it 




r 








•/ 


< 
< 


S 










y 










/ 




y 










J y 


Line 






Running winding 
and poles 






\ 


' 



Figure 6.64.— Starting and running stator windings. 



Starting 




Centrifugal switch 
closed on start 



B Running 
I jj Running 

II y 

Line j windin ^ 
, J Starting 

Switch opens at 
75% of speed 



Figure 6.65.— Centrifugal switch to remove starting winding. 



Centrifugal switch 



Capacitor 



Running Starting 

wind^ S winding 



Figure 6.66.— Capacitor-start motor. 



158 



REFERENCES 

1. Allis-Chalmers (Milwaukee, WI). Motor Control-Theory and 
Practice. 1955. 

2. Bergmann, R. W. Excavating Machinery. Ch. in Standard 
Handbook for Electrical Engineers. McGraw-Hill, 10th ed., 1968. 

3. Fitzgerald, A. E., C. Kingsley, Jr., and A. Kusko. Electric 
Machinery. McGraw-Hill, 3d Ed. 1971. 

4. Hardie, R. C. Mine Hoists. Ch. in Standard Handbook for 
Electrical Engineers. McGraw-Hill, 10th ed., 1968. 

5. Hugus, F. R., J. A. Buss, and E. L. Parker. Mining Machine 
Motor Characteristics. Min. Congr. J., v. 41, May 1955. 

6. Mining Machine Motor Identification. Min. Congr. J., 

v. 41, Mar. 1955. 

7. Joy Manufacturing Co. (Franklin, PA). Direct Current Min- 
ing Machinery. 5th ed., 1971. 

8. Kosow, I. L. Electric Machinery and Transformers. Prentice- 
Hall, 1972. 

9. Lloyd, T. C. Electric Motors and Their Applications. Wiley- 
Interscience, 1969. 



10. Marion Manufacturing Co. (Marion, OH). 191-M Mining 
Shovel. Doc. Specification 542-5, 1979. 

11. Marklekos, V. E. Electric Machine Theory for Power 
Engineers. Harper and Row, 1980. 

12. Millermaster, R. A. Harwood's Control of Electric Motors. 
Wiley-Interscience, 4th ed., 1980. 

13. Morley, L. A. Utilization and Efficiency in Underground 
Coal Mine Electrical Systems. Paper in Mine Power Distribution. 
Proceedings: Bureau of Mines Technology Transfer Seminar, Pitts- 
burgh, Pa., March 19, 1975. BuMines IC 8694, 1975. 

14. Oscarson, G. L. The ABC's of Large Induction Motors. E-M 
Synchronizer. Electrical Machinery Manufacturing Co., Min- 
neapolis, MN, 1955. 

15. Smeaton, R. W. (ed.). Motor Application and Maintenance 
Handbook. McGraw-Hill, 1969. 

16. Stefanko, R. Coal Mining Technology Theory and Practice. 
Soc. Min. Eng. AIME, 1983. 

17. Tsivitse, P. J. Mining Motors. Ch. in Motor Application and 
Maintenance Handbook, ed. by R. W. Smeaton. McGraw-Hill, 1969. 



159 



CHAPTER 7.— GROUNDING 



A vital part of any mine power distribution system is 
the connection to earth or ground, which is referred to as 
the mine grounding system. It consists of grounded or 
grounding conductors, extending from ground beds to 
equipment. A grounded conductor is a power conductor 
tied to the grounding system; a grounding conductor is 
separate from the power conductors and is used only to 
ground exposed metallic parts of the power system. A 
ground bed, also termed a ground mesh or grounding 
electrode, as well as other names, is a complex of conduc- 
tors placed in the earth to provide a low-resistance con- 
nection to "infinite" earth. The grounding system serves 
to protect personnel and machinery from the hazards 
associated with electrical equipment that is operating 
improperly. The protection afforded can be divided into the 
following four functions, which are the main purposes 
behind grounding the system. 

First, the grounding system must limit potential 
gradients between conducting materials in a given area 
(38). 2 During a ground fault, for instance, a phase conduc- 
tor comes into contact with a machine frame, and current 
flows through the equipment; subsequently, the potential 
of the equipment tends to become elevated above ground 
potential by an amount equal to the voltage on the 
conductor. If a person touches the machine, while being 
simultaneously connected to ground in some manner, the 
body's potential can become elevated, possibly to a lethal 
extent. The maximum potential to which a person could be 
exposed when touching a machine frame is equal to the 
voltage drop along the grounding conductors. Thus, the 
grounding system must provide a low-resistance path for 
the fault current to return to the source, and the ground 
conductors should have low resistance so they can carry 
the maximum expected fault current without excessive 
voltage drop. An example of the exposed potential in a 
surface mining situation is illustrated in figure 7.1 (38). 

Second, the grounding system should limit the energy 
available at the fault location. Heavy arcing or sparking 
can ignite nearby combustible material. The air itself can 
become ionized, making it capable of carrying tremendous 
amounts of current. A high-energy fault can vaporize 
breakers, switchgear, and phase conductors, and protec- 
tive enclosures may be blown apart with explosive force 
(21). Controlling the maximum allowable fault current 
significantly reduces the danger of fire and holds equip- 
ment damage to a minimum. 

Third, the control of overvoltages is essential. An 
overvoltage condition may occur by accidental contact of 
equipment with a higher voltage system, or from transient 
phenomena due to lightning strokes, intermittent ground 
faults, autotransformer connections, or switching surges 
(4). The maximum ratings for cable insulation, trans- 
former windings, relay contactors, and so forth may be 
temporarily exceeded in these cases. This does not usually 
result in an immediate breakdown of equipment, but 
component parts of the electrical system are successively 
overstressed and weakened by repeated exposure (see 



1 The author wishes to thank Alan M. Christman, who prepared the 
original material for many sections of this chapter while he was a graduate 
student at The Pennsylvania State University. 

2 Italicized numbers in parentheses refer to items in the list of references 
at the end of this chapter. 



chapter 11). This leads to premature failures, reduced 
component life, and mysterious "nuisance trips," which 
can occur without apparent reason. By providing a path 
between the transformer neutral and ground, most of the 
sources of transient overvoltages can be reduced or possi- 
bly eliminated. 

Last, a grounding system should isolate offending 
sections by selective relaying of ground faults (44). The 
sensitivity and time delays of the protective circuitry 
should be adjusted so a fault in a certain area will cause 
the local breaker to sense the malfunction and quickly 
remove power from only the affected section. If the relative 
tripping levels and speeds are not established correctly, 
nearby breakers may not trip when they should, and a 
small problem could escalate into a large calamity. Con- 
sequently, power to half a mine may go out because of poor 
relay coordination, and much time could be lost in the 
effort to trace and locate the trouble spot. Thus, the 
relaying system must be arranged so, even at the lowest 
level of the power-distribution chain, sufficient fault cur- 
rent can flow to enable the protective circuitry to sense it 
and take remedial action. 

Chapters 9 and 10 cover the protective circuitry used 
to provide the function of section isolation, while chapter 
11 describes the devices employed with the grounding 




3- phase diagram 

Phase-conductor resistance 
-AAM VWV — 



31 



Neutral 
resistor 



Total 

fault 

current 



Current through 
grounding conductor 



. Total fault 
Line-to-ground j current 
fault on shovel 

^* f Shovel frame is at 
this potential 



^VWV 



-WJV- 



1 Grounding-conductor 

resistance 
Safety ground-bed R s 

resistance T . 

Current through earth 



_ -2^- Potential to 
A which operator 
may be 
p § subjected 

U Im 



Circuit for line-to-ground fault 
Figure 7.1.— Illustration of electrical shock hazard. 



160 



system for transient overvoltage control. As an introduc- 
tion, this chapter looks mainly at the first three purposes 
and presents the common methods of system grounding, 
the effect of electric shock on human beings, mine ground 
system characteristics, and ground-bed construction. Ex- 
tensive information about grounding is contained in prac- 
tically all subsequent chapters. 



GROUNDING SYSTEMS 



figure 7.3, now has its neutral solidly referenced to ground 
(20). The hazards of this system are due to the magnitude 
of the fault current. Detection equipment must be sensi- 
tive enough to detect low-level fault currents and fast 
enough to disconnect bad circuits before heavy faults can 
disrupt system integrity. Large fault currents, typically 
several thousand amperes, can explode protective enclo- 
sures, destroy equipment, and start fires, which is an 
excellent reason for not using this technique in explosive 
atmospheres. 



Over the past few decades, several different grounding 
philosophies have held sway in the electrical industry, 
each with its own advantages and disadvantages (20). 
These methods of grounding are discussed below. Note 
that reactance-grounded systems are not presented in the 
following paragraphs, as they are not normally used in 
industrial power systems. 

Ungrounded Neutral 

The ungrounded system was probably the first to be 
used because of its simplicity. Here there are no inten- 
tional ground connections in the system whatsoever. How- 
ever, a perfect ungrounded system cannot exist, since any 
current-carrying conductor may be coupled to ground 
through numerous paths, including the distributed capac- 
itance of its wiring, or through motor windings (49). This 
phenomenon is shown in figure 7.2 (20). The first line- 
to-ground fault on such a system will have very little effect 
(27) because there is no way for the fault current to find a 
complete circuit back to the source, and its magnitude will 
be very small or nil. Very low fault current means no flash 
hazard and no equipment damage. Circuit operation con- 
tinues normally with no interruption of power, an impor- 
tant consideration in industries where downtime is criti- 
cal (60). The first fault is often hard to locate because its 
effects are negligible. Often no repair effort is made until 
a second fault occurs, with its concomitant hazards of 
arcing, heavy current flow, and equipment damage. Since 
the entire system is "floating," there is no control of 
transient overvoltages. Except for the problem of acciden- 
tal contact with a higher voltage system, all the other 
overvoltage sources mentioned previously are enhanced 
because of distributed capacitance to ground (20). 

Solidly Grounded Neutral 

An alternative is the solidly grounded neutral (20). 
The first ground fault produces a substantial neutral 
current flow, which may be quickly sensed by protective 
circuitry, thereby shutting down the bad section. Overvolt- 
ages are controlled since the system, as illustrated in 



Supply transformer 




Ground 



"'^Ja^^ balanced 
Figure 7.2.— Capacitance coupling in ungrounded system. 



Low-Resistance Grounded Neutral 

The low-resistance grounded-neutral system is estab- 
lished by inserting a resistor between the system neutral 
and ground. The resistance is such that ground-fault 
currents are limited from 50 to 600 A, but are commonly 
about 400 A (20). Transients are controlled by the ground 
connection, and ample fault current is available for actu- 
ating protective relays. The flash hazard is not as serious 
as in the solidly grounded neutral system, but a current 
flow of 400 A can still do considerable damage. To limit 
damage, the least sensitive ground relay should respond to 
10% of maximum ground-fault current. A schematic dia- 
gram of this method is shown in figure 7.4 (20). 

High-Resistance Grounded Neutral 

Perhaps the best technique, and that required by law 
in coal-mining applications on portable or mobile equip- 
ment, is the high-resistance grounded system, often re- 
ferred to as the safety ground system. The neutral ground- 
ing resistor is sized according to the system voltage level, 



Supply 
transformer 
secondary 




Ground 
Figure 7.3.— Solidly grounded system. 



Supply 
transformer 
secondary 




Figure 7.4.— Resistance-grounded system. 



161 



in general to limit ground-fault current at 50 A or less. 
Where the line-to-neutral potential is 1,000 V or less, the 
grounding resistor must limit fault current to 25 A or less; 
above 1,000 V, the voltage drop in the grounding circuit 
external to the resistor must be 100 V or less under fault 
conditions, With this system, sensitive relaying must 
detect faults on the order of a few amperes to provide fault 
isolation and facilitate quick location of the trouble spot 
(60). The level of fault current is also low enough to 
practically eliminate arcing and flashover dangers. The 
ground connection also serves to limit the amplitude of 
overvoltages. However, loads cannot be connected line to 
neutral, as the grounding conductor must not carry any 
load current. 



ELECTRIC SHOCK 

For a safe grounding system to be efficiently and 
economically designed, voltage and current levels that are 
harmful to human beings must be determined. With the 
trend toward larger and more powerful mining machinery, 
distribution voltage and current levels have risen propor- 
tionately. Constant vigilance is required when using elec- 
tricity if the hazard of electrocution is to be avoided. Even 
if a shock is nonlethal, involuntary movement caused by 
the shock may lead to serious injury or death. As an 
example, a man standing upon a ladder may come into 
contact with a live wire and fall from his perch (12). 

Physiologically speaking, the muscles of the body are 
controlled by electrical impulses transmitted from the 
brain via the nervous system. These pulses occur at a rate 
of about 100 per second and may be of positive or negative 
polarity. From this, it can be seen that the human "in- 
ternal power supply" operates at about 50 Hz, which is 
exactly the frequency of the electric power generated in 
Europe, and is only 10 Hz removed from the U.S. power 
generation frequency of 60 Hz. This is an unfortunate 
coincidence, for tests have shown that the most dangerous 
frequencies to which a person can be exposed are power 
frequencies in the range of 50 to 60 Hz (12). 

How sensitive are human beings to the flow of elec- 
tricity? Tests have indicated that for an average male 
holding No. 7 AWG (American Wire Gauge) copper-wire 
electrodes in his hands, 60-Hz ac is first perceived at a 
level of about 1 mA (12). By intermittently touching or 
tapping an electric conductor, currents of only 1/3 mA can 
be felt. In the case of dc, the threshold of perception for the 
average male is 5.2 mA. Sensitivity levels for women in 
the cases mentioned above can be found by multiplying 
the male values by a factor of two-thirds (13). It is gener- 
ally agreed that the magnitude and duration of the cur- 
rent are the important shock parameters, rather than the 



potential difference or voltage (12), as can be seen in table 
7.1 (42). 

As current magnitude is increased above the level of 
perception, many test subjects have reported a tingling 
sensation, the intensity of which increases as the current 
rises. Generally, muscles in the vicinity of the current 
path start to contract involuntarily, until finally a point is 
reached where the subject being tested can no longer 
release his grip on the conductor (14). The maximum 
current magnitude that a person can withstand while still 
able to release the live conductor through the use of 
muscles stimulated directly by the current, is called the 
let-go current (fig. 7.5) (14, 16). Tests performed on hun- 
dreds of volunteers have shown that the maximum let-go 
current for a healthy adult male is 9.0-mA ac and 60-mA 
dc. The corresponding values for women are 6.0-mA ac and 
41-mA dc. These safe-limit values apply to 99.5% of the 
sample population (11). The value of a specific individual's 
let-go current is virtually constant, even with repeated 
exposures to that current level. In addition, these multiple 
exposures can be tolerated with no ill effects (16). 

It has been stated that human tissue possesses a 
negative resistance characteristic. In other words, an 
increase in current magnitude or contact duration leads to 
a decrease in the value of skin resistance (17). In any case, 
if a person has grasped a live conductor and realizes that 
he/she cannot let go, fear-induced perspiration will cause a 
lowering of the body's resistance, and more current will 
flow. For ac, when the current level across the chest 
reaches more than 18 to 22 mA, the chest muscles tighten 
involuntarily and breathing ceases. Although circulation 
of blood by the heart is unimpaired, death by asphyxiation 
can occur within minutes (43). 

If an individual's initial contact with a live wire 
causes a current flow ranging from about 50 to 500 mA, 
ventricular fibrillation may result (48). Under normal 
conditions, the heart beats with a strong, coordinated 
rhythm. However, a current passing through the heart 
when the ventricles (the heart's two large pumping cham- 
bers) are just starting to relax after a contraction, can 
cause the various fibers of the heart muscle to beat weakly 
in an uncoordinated manner (43). In this condition, known 
as ventricular fibrillation, the heart is almost totally 
incapacitated and blood circulation decreases practically 
to nothing. Within 2 min, the brain begins to die because 
of oxygen deficiency. Once initiated, ventricular fibrilla- 
tion almost never stops spontaneously, and treatment by 
trained medical personnel must be secured if the victim is 
to survive. 

Obviously, people cannot be used as test subjects in 
ventricular fibrillation experiments because of the high 
risk involved. Numerous tests have been carried out on 
several species of animals and the results extrapolated, 



Table 7.1. —Current range and effect on a typical man weighing 150 lb 



Current 

Less than 1 mA.. 
1 mA 

3 mA 

10 mA 

30 mA 

75 mA 

4 A 

Greater than 5 A. 



Physiological phenomena Effect on man 

None Imperceptible. 

Perception threshold Mild sensation. 

Pain threshold Painful sensation. 

Paralysis threshold of arms and hands Person cannot release hand grip; if no grip, victim may be thrown clear. 

Tighter grip because of paralysis may allow more current to flow; may be 

fatal. 

Respiratory paralysis Stoppage of breathing, frequently fatal. 

Fibrillation threshold (depends on time) Heart action uncoordinated, probably fatal. 

Heart paralysis threshold (no fibrillation) Heart stops on current passage, normally restarts when current interrupted. 

Tissue burning Not fatal unless vital organs are burned. 



162 



Ld 

X 
o 

9 
I 



30 r 



20 



10 




_L 



J L 



J 



5 10 50 100 5001,000 5,000 

FREQUENCY, Hz 

Figure 7.5.— Effect of frequency on let-go current for men. 



based upon body weight, to cover human beings (15). It was 
found that fibrillating current is proportional to body 
weight and inversely proportional to the square root of the 
shock duration. Using 50 kg (110 lb) as a body weight, it 
has been proposed that the value of current that can be 
safely endured by 99.5% of normal adults without causing 
ventricular fibrillation is (16) 



I = 



116 



5s 
8.3 ms 



(7.1) 



where I = rms ac, mA, 
and t = shock duration, s. 

As noted, this equation is valid for values of time between 
8.3 ms and 5.0 s (15). 

It may be seen from the above equation that for a 1-s 
contact time, the ventricular fibrillation threshold current 
is about 116 mA. Since a normal person has a pulse rate 
between 60 and 80 beats per minute, the critical phase of 
the heartbeat (when a person is vulnerable to ventricular 
fibrillation) occurs about once each second. Therefore 
during a shock lasting for 1 s or more, the heart must pass 
through this critical phase (48). As a result, it is thought 
that ventricular fibrillation is the leading cause of death 
by electric shock. 

Higher currents on the order of a few amperes will 
freeze both the chest and heart muscles, thereby prevent- 
ing the onset of ventricular fibrillation. Generally, the 
heart will restart upon the cessation of current flow (48). 
These current magnitudes are less dangerous statistically 
than the lower values where fibrillation is prevalent. 
Further increases in current level, to 5A and above, may 
produce serious burns leading to shock and possible death, 
while current levels that substantially elevate body tem- 
perature produce immediate death (16). 

In an electric-shock situation, the victim's electrical 
resistance plays an important role in determining how 
much current will flow, as indicated by Ohm's law: 



I = 



(7.2) 



For a human being, at least three components of resis- 
tance have been isolated: contact resistance, skin resis- 
tance, and internal resistance (43). Contact resistance, as 



illustrated by table 7.2, depends upon the degree of skin 
moistness and the area of contact with the live conductor 
(42). Values of 40,000 to 50,000 fl/cm 2 are given for dry 
skin and 1,000 ft/cm 2 for wet skin (13). Skin resistance 
depends upon the physical condition of the tissues: A 
person who does rough, heavy outdoor work may have a 
skin resistance of 10,000 fi, while a value of 1,000 Q is 
typical of a sedentary office worker (43). Internal resis- 
tance is the resistance of the body's interior and is gener- 
ally accepted to be about 500 ft between major extremities 
(25). 

Table 7.2.— Typical resistance for various contact situations, 
ohms 

Contact 1 Dry skin Wet skin 

Finger touch 500,000 20,000 

Hand on wire 50,000 10,000 

Finger-thumb grasp 20,000 5,000 

Hand holding pliers 20,000 2,000 

Palm touch 10,000 1,000 

Hand holding 1.5-in pipe 2,000 500 

2 hands holding 1.5-in pipe 500 NA 

Hand, immersed NAp 200 

Foot, immersed NAp 100 

NA Not available. NAp Not applicable. 

1 Skin surface only; resistance may be lower when skin is cut, blistered, or 
abraded. 

Voltage magnitude has some effect upon the body's 
reaction to electric shock, although current is by far the 
most important parameter. Potentials greater than about 
240 V simply puncture the skin, thereby negating the 
effects of skin resistance (12). There is also some evidence 
that overall body resistance varies inversely with the 
applied voltage, although this is subject to disagreement. 
The relationship is given by (43). 



R oc E 



(7.3) 



where R = resistance, 12, 
E = potential, V, 
and n = 1.5 to 1.9. 

Above about 2,400 V, tissue damage due to burning 
becomes the major cause of electric-shock injury (42). 

Thus it can be seen that the body's response to 
electricity is extremely complex, and currents on the order 
of a few milliamperes can be fatal if long continued. 

CHARACTERISTICS OF MINE GROUNDING 
SYSTEMS 

The concept of protecting mine electrical equipment 
and personnel against the consequences of ground faults 
by suitable grounding has existed since electricity was 
first introduced into coal mines. As early as 1916, the 
Bureau of Mines recommended equipment frame ground- 
ing as a means of preventing electrical shock to miners 
working on or around electrical equipment (6). For the coal 
mining industry, a suitable grounding system has always 
been a difficult problem, more complex and difficult than 
in other industries. 

Ground Beds 

For mine usage, the electrical distribution cables and 
overhead transmission circuits carry into the mine one or 



163 



more grounding conductors in addition to the phase con- 
ductors. Each piece of ac equipment has its frame solidly 
connected via these grounding conductors to a safety 
ground bed commonly located near the main surface 
substation and consisting of buried horizontal conductors 
or driven rods, or a combination of both. The neutral of the 
substation transformer secondary is also connected to the 
safety ground bed through the neutral grounding resistor, 
as shown in figure 7.6. It should be noted that many 
important components are missing from this diagram, and 
chapter 13 covers substation circuitry in detail. 

The substation actually requires two ground beds, 
maintained some distance apart. Lightning discharges 
and other transformer primary surging conditions are 
directed to the system or station ground. The system and 
safety grounds must be kept separate so current flow 
intended for one will not enter the other. It is essential for 
the safe operation of the mine power system that the 
resistance of the beds be maintained at 5.0 U or less (3, 39, 
44). A ground bed with this resistance range is often 
termed a low-resistance ground bed. 

To demonstrate one reason for a low-resistance bed, 
consider a situation where lightning strikes the substa- 
tion, and 10,000 A is discharged through the surge arrest- 
ers into the system ground bed. If the ground bed is of 5.0-fi 
resistance, a potential of 50,000 V is developed, and the 
grounding grid of the ground bed becomes elevated to 50 
kV above infinite earth. Depending upon the physical 
extent of the grid, a person walking through the area 
underlain by the grid could bridge a lethal potential 
gradient with his or her feet (2). Metallic objects within 
the potential gradient field can also be elevated to danger- 
ous potentials and become lethal to the touch. Typical 
step and touch potentials are illustrated in figures 7.7 and 
7.8 (2). 

These step and touch potential hazards are applicable 
to both the system and safety ground beds. However, the 
dangers of a high-resistance safety ground bed are not 
found close to the bed but at the mining equipment. The 
most insidious feature of the safety ground system is that 
the equipment connected to it is maintained not at earth 
potential, but at the safety ground-bed potential. Unless 
the bed has low resistance, any safety ground-bed current 
flow can render every piece of mine equipment potentially 
lethal. The flow can be created by faults to earth, coupling 
from lightning strokes to the system ground, lightning 
strokes to safety grounded machinery, and stray currents 
from dc haulage systems. Three such cases are illustrated 
in figures 7.9, 7.10, and 7.11 (9). Consequently, with 
high-resistance ground beds, an elevated frame potential 
is a problem not just on the machine where it occurs, but 
everywhere (10). 




Rf? $ f 
"flfc 

1 — wwwwwww 



IR 



1 "2 



Figure 7.7.— Step potentials near grounded structure. 




Potential rise above 
remote earth during fault 



0.5 R f 

I A/WVWWVWV 1 



Figure 7.8.— Touch potentials near grounded structure. 



Power center 



Load 



Incoming 
power 



> 



Surge 
arresters 9 



System ^. 
ground = 



Substation 
transformer 



Neutral 
resistor 



-± Safety 
- ground 



3-phase 
power to 
equipment 

Grounding 
conductor 



Figure 7.6.— Simplified one-line diagram of substation. 




to machine frame 
W Safety ground bed 



ys Fault to 



Figure 7.9.— Line-to-earth fault resulting in current flow 
through safety ground bed. 



164 



Substation 




Figure 7.10.— Lightning stroke to equipment causing current 
flow through safety ground bed. 




Mining 
machine 



Safety 

ground 

bed 



Figure 7.11.— Lightning stroke current through system 
ground bed causing elevation of safety ground bed. 



Grounding in Underground Mining 

Early practice in underground coal mining was to 
drive a metal rod into the mine floor and use that as a 
ground. In almost every case this arrangement proved to 
be totally unacceptable, with test measurements indicat- 
ing 25-fi or more resistance (28). With the exception of 
pumps, the contact resistance of mining machinery with 
the mine floor also proved to be too high for adequate 
grounding. Rail haulage track systems, even though often 
poorly bonded, showed much lower resistance to ground 
than most metallic rods driven specifically for that pur- 
pose. As a solution, Griffith and Gleim (28) in 1943 stated 
that ". . . consideration should be given to a grounding 
circuit carried to the outside of the mine." Present coal 
mine practice does just that. 

A simple form of the bipower (mixed ac-dc) system in 
use in underground coal mines today is illustrated in 
figure 7.12. After transformation, three-phase ac power 
enters the mine to supply the various three-phase ac loads. 
Some of the ac power is converted to dc at rectifier stations 
to power the locomotive system and, occasionally, dc face 
equipment. More often, any dc face machinery is powered 
from rectifiers located in the mine section. Except for the 
trolley system, all dc as well as the ac equipment frames 
are connected to a common junction, which is tied to the 
surface safety ground bed. In order for the system to be 
effective, grounding conductors must be continuous and 




ac 

loads 



Figure 7.12.— One-line diagram of simplified mine power 
system. 



this continuity must be verified. Ground-check monitors 
ensure this. 

Trolley locomotives generally utilize the overhead 
trolley wires as the positive conductor and the tracks as 
the negative. Neither of these is tied to the rectifier- 
station frame ground. However, because the track is in 
contact with the mine floor, the negative conductor for the 
trolley system is grounded. The dc system that supplies 
power to face equipment normally employs trailing cables 
that have neither the negative nor positive conductor 
grounded. Thus, this subsystem is often ungrounded un- 
less the supply is obtained from the trolley system. Note 
that in diode-grounded systems, the negative conductor is 
grounded. 

At each transformation step within the power system, 
such as in a power center, an additional neutral point must 
be established on the transformer secondary. The neutral 
is tied through a grounding resistor to the equipment 
frame and thus via the grounding conductors to the safety 
ground bed (an exception will be discussed later). 

Even with all these grounding points, the ac ground- 
ing system must be isolated from separate dc power 
systems. If it is not, dc may appear in the ac grounding 
system, thus elevating it above true ground potential. If an 
ac ground current is present, it will be offset by the dc 
level. The principal concern is with trolley installations, 
where isolation is achieved by having no common points 
between the ac and dc systems. Various techniques have 
been tried to maintain separation or to eliminate dc offsets 
while grounding dc face equipment frames. 



165 



Face Equipment Grounding 

When a working section utilizes an ac continuous 
miner energized from a section power center and dc 
shuttle cars powered from the trolley system, the ground 
potentials of the dc and ac equipment frames are not 
necessarily equal, because of the voltage drop in the track. 
Jacot (33) suggested that this problem could be solved by 
isolating the low-voltage ac neutral point from the power- 
center frame and also the high-voltage grounding system, 
and connecting it via an insulated cable to the track, as 
shown in figure 7.13. The low-voltage neutral point re- 
mains connected to the ac face-equipment frames. This 
technique should make the low-voltage ac and dc equip- 
ment frame potentials the same, thus eliminating dc offset 
problems. Difficulties can still arise with this method. If 
any track rail bonds are bad between the ac and dc 
low-voltage ground points, the dc frame potentials might 
be elevated with respect to the ac frames. Further, the 
power center must be constantly maintained at a safe 
distance from the tracks to preserve isolation between the 
track and high-voltage grounding systems. 

Another method is shown in figure 7.14A. Here, a 
section power center supplies power to ac face equipment 
and also, through a rectifier, to dc machinery (usually 
shuttle cars). The rectifier is isolated within the section 
power center, but the dc output is grounded through a 
center-tapped current-limiting resistor. All dc equipment 
frames are then grounded by trailing-cable grounding 
conductors, which in turn are connected to the center tap 
of the grounding resistor. The latter point is connected to 
the high-voltage grounding system. This has been consid- 
ered a very safe dc ground protective system because it 
permits the use of protective circuitry to trip the rectifier 
breaker in case of a dc ground fault (see chapter 9). 
However, the use of the center-tapped resistor has been 
criticized (46). On such a system, any failure to maintain 
grounding-conductor conductivity or accidental connec- 
tion of a wrong conductor when splicing cables may lead to 
a hazard. Nevertheless, an important advantage of the 
method is that the dc and ac frame potentials can be the 
same. A more recent method for limiting dc ground-fault 



current is similar to high-resistance ac grounding and is 
illustrated in figure 7.145. 

In 1963 the Bureau of Mines accepted the use of 
silicon diodes as a means of grounding dc face equipment 
frames. When a diode is used, the grounding resistors are 
not needed because the frame is grounded through the 
diode to the negative conductor, as illustrated in figure 
7.15. The diode circuit also includes a ground protective 
device, which will interrupt the power if a current flows 
from a positive power conductor to an equipment frame 
(again, see chapter 9). According to Jones (37), diode 
grounding should ensure good ground continuity since the 
same conductor acts as both a dc negative conductor and 
the grounding connection. However, a grounding diode 
only protects the dc system against ground faults within 
the equipment frame. Current leakage to ground or faults 
within trailing cables can still present hazards. 




A Resistors between dc line conductors and grounding conductor 



Grounding 
resistor 




From safety 
ground bed 



Ground conductor from 
power-center frame to 
safety ground bed 

Power center 



Continuous 
miner(ac) 



J7 



Neutral ^ 
resistor § 



Lifted from ' 
frame 



Insulated grounding 
<s conductor 



+ - 



M 



Shuttle 
car (dc) 



Grounding 
conductor 



Trolley system 



Rail in contact with mine floor 

Figure 7.13.— Mixed ac-dc mine power system; dc load 
energized from trolley system. 



B Resistor between transformer neutral and grounding conductor 

Figure 7.14.— System grounding with current-limiting 
resistors. 



Load center 



To 
rectifier 



Machine frame 



Connected to frame 
From safety ground bed through diode 



CB 



) 



Grounding diode 
Figure 7.15.— Diode grounding of machine frame. 



166 



Track Grounding 

As previously mentioned for trolley systems, the rec- 
tifier frame is grounded by the ac system, but the negative 
conductor is grounded to the mine floor through the track. 
In order to maintain isolation, there is no internal connec- 
tion between the rectifier output (or the trolley distribu- 
tion system) and its frame. However, if the rectifier is 
sitting on the mine floor, there is a possible common point 
from the track (dc) to the rectifier frame (ac). Ideally, the 
common point through the earth is a much higher resis- 
tance than the rail itself so that all rectifier current 
returns in the rail. When the rail resistance increases 
because of poor bonding or crossbonding, some current 
may flow through the earth to or from the rectifier frame, 
depending on the rail potential. Thus dc is introduced in 
the ac ground system. 

Leakage of trolley- wire insulator to the roof or rib may 
have the same effect, although it is less common. This lack 
of effective separation can cause dc offset currents on any 
mining machine and electrical system whenever the sum 
of the mine floor resistance and equipment frame contact 
resistances is too low and, therefore, dc current flow is 
permitted through the earth. To help minimize any prob- 
lems, rectifiers should be located no closer than 25 ft from 
the track. In severe cases, the rectifier frame can be 
insulated from the mine floor. 

The preceding has shown that haulage conversion 
units are the primary source of dc offset currents. Regard- 
less of the source, once stray dc currents occur, they can 
exist on all the ac grounds within the mine. This problem 
is further complicated since these currents may also travel 
through water pipes and hoses, or anything conductive. 

The two most undesirable effects of dc offset currents 
on the ac ground system are nuisance tripping and inter- 
machine arcing. Nuisance tripping can occur whenever 
the offset ac waveform is greater than the relay trip value, 
and it primarily affects ground-overcurrent relays and 
ground-check monitors. Intermachine arcing occurs when 
two machine frame potentials are not the same. While 
they are touching, a current flow is possible, but when 
they separate, arcing may occur. These problems are 
discussed further in chapter 17. 

Grounding in Surface Mines 

The typical grounding system for a surface coal mine 
is similar to that for underground mining. One or more 
substations with resistance-grounded secondaries are em- 
ployed to transform the incoming utility voltage to the 
lower potential used by the mining machines. At this 
level, pit distribution is carried on overhead lines or cables 
to supply switchhouses located near the particular piece of 
equipment. A trailing cable completes the power circuit 
from the switchhouse to the machine. A switchhouse is 
sometimes connected via cable to a portable substation, 
which supplies lower voltage power to production, auxil- 
iary, or lighting equipment. 

Substation grounding includes both a system and a 
safety ground bed, each physically removed and electri- 
cally isolated from the other. Grounding conductors extend 
from the safety ground bed to all equipment frames. The 
neutrals of the transformer secondary of portable substa- 
tions are resistance grounded to the equipment frame. 

In contrast with underground coal mines where the 
entire secondary distribution system is underground, both 



the primary and secondary lines in a surface mine are out 
in the open where they are exposed to lightning. In fact, 
equipment such as draglines and shovels are subject to 
direct strokes (fig. 7.10). For the best possible protection 
from lightning, it is essential that the grounding system 
have as low a surge impedance as possible. The key factor 
here is to provide many short, direct paths to earth. The 
specifics of lightning protection for all mines are presented 
in chapter 11. 



GROUND-BED CONSTRUCTION 

Since the minesite is determined by the location of the 
rock or mineral to be extracted, the conditions required for 
the installation of an adequate ground bed are not always 
easily met. If annual rainfall is low or soil resistivity is 
high, an extensive array of buried metallic conductors may 
be necessary to assure a low-resistance connection to 
earth. Measurement of soil parameters can be made 
before the construction of a grounding grid is begun, 
thereby ascertaining the configuration for the metallic 
network that will yield the desired values of earth resis- 
tance and potential gradient. After construction, the re- 
sistance of the selected ground-bed configuration must be 
checked. Proper design at the time the ground bed is 
installed will save much time and expense in later years. 

Present-day ground beds can be divided into two 
general categories: meshes and rodbeds. A mesh is a 
horizontal network of metallic conductors arranged in a 
grid pattern, which is embedded a short distance below 
the earth's surface. A rodbed is an interconnected network 
of vertical metallic rods driven into the earth. The metal- 
lic components for either ground-bed type are also called 
electrodes. 

Ground Resistance 

Any grounding system exhibits some finite resistance 
with respect to infinite earth, even though it is completely 
immersed in the soil. When a fault from a power conductor 
to earth occurs, current can flow through the ground-bed 
metallic electrodes, across the soil-metal interface, and 
into the ground. The greater the surface area of metal in 
intimate contact with the soil, the lower the resistance. 
Most of the actual resistance exhibited by each metallic 
conductor occurs within 6 to 10 ft of the electrode, as 
illustrated in figure 7.16 (36). If the surrounding soil is 
viewed as a succession of concentric shells, it is easily seen 
that the shells adjacent to the electrode have a much 
smaller cross-sectional area, and hence a higher resistance 



r 



Grounding 
electrode 



TT TTTT T 
! mm 



III"! 



iii: !!iv u ;!Jj M iJ 



Figure 7.16.— Resistance of earth surrounding electrode. 



167 



than more distant shells. Consequently, the main factors 
which determine grounding-grid resistance are the physi- 
cal dimensions of the system and the innate characteris- 
tics of the soil, primarily its resistivity (51). Figure 7.17 
shows how the total resistance of a driven rod varies as it 
penetrates soil horizons of different resistivity (36). 

The electric field around a current-carrying wire is 
analogous to the electrostatic field surrounding a charged 
conductor of similar shape. By calculating the capacitance 
of an electrode immersed in the soil, its resistance can 
then be determined. For a conductor buried deeply in the 
earth (29), 



R = 



P 
4ttC 



(7.4) 



where R = resistance, Q, 

p = soil resistivity, Q-m, 
and C = capacitance, F. 

If the conductor is relatively near the earth's surface, as is 
usually the case, the effects of the conductor image, which 
is located an equal distance above the surface, must be 
included in the formula, yielding (29) 



R = 



P 
2ttC- 



(7.5) 



is ineffective. However the conductance curve is almost 
linear. Figure 7.19, which is an extended version of the 
previous graph, shows very clearly that even at depths of 
100 ft, the conductance increases in direct proportion to 
the length (23). If the soil can be easily penetrated, deeper 





5 
10 
15 
20- 
25 



Surface of earth 



\ 



_L 



350 250 150 

RESISTANCE, XL 



50 



' i ' 

Light | \ 
sand ; 






Hard sand ; 


- 

Clay mixed 
with sand 












r 


, 



Test boring 



Figure 7.17.— Decrease in earth resistance as electrode 
penetrates deeper soil horizons. 



For multielectrode systems, the capacitance of each 
conductor plus its mutual capacitance with respect to all 
other conductors must be calculated. By maximizing the 
capacitance, the resistance can be minimized, which is the 
desired goal. 

The two predominant methods for determining the 
capacitance of earth-electrode systems are Howe's average 
potential method and Maxwell's method of subareas, each 
of which has a constant charge density and potential (29). 
Howe's technique assumes a uniform charge density on 
each electrode and then calculates the average surface 
potential. The capacitance can be found from (21) 



C = 



(7.6) 



where C = capacitance, F, 

Q = charge, C, 
and V = potential, V. 




10 15 

DEPTH OF ROD, ft 



0.06 



.02 



in 

u 

-z. 
< 

o 

Q 

o 
o 



25 



Figure 7.18.— Calculated values of resistance and conduc- 
tance for 3 /4 -in rod driven to depth of 25 ft. 



Electrode Configuration Formulas 

One of the most common ground systems is the 
rodbed. For a single vertical rod (22), 

R = 2^0nf-1), (7.7) 

where p = soil resistivity (ft-m or fi-ft) 

I = length of rod, m or ft, 
and a = radius of rod, m or ft. 

Figure 7.18 shows how the resistance and conductance of 
a typical driven rod vary as the rod length is increased 
(23). It can be seen that the resistance curve starts to 
flatten out, which indicates that a length in excess of 15 ft 




40 60 

DEPTH OF ROD, ft 

Figure 7.19.— Calculated values of resistance and conduct- 
ance for Vi-in rod driven to depth of 100 ft. 



168 



rods are always better. The simple nomogram shown in 
figure 7.20 may be used to estimate the resistance of a 
driven rod without carrying out calculations (55). Figure 
7.21 shows the effect of soil resistivity on the resistance of 
a driven rod, as well as the benefits gained by using longer 
rods (22). It could be pointed out that several shorter 
ground rods are easier to drive than one long rod of the 
same total length. However, when using multiple rods, the 
effects of mutual resistance tend to negate some effective- 
ness, so the resistance of the group is greater than would 
be expected unless very large spacings are used between 
electrodes (23). Figure 7.22 shows this effect for rods 
spaced at a distance equal to their length, while figure 
7.23 shows the advantages that accrue when spacing is 
increased from 0.5 to 100 ft (54). 



One of the following two formulas can be applied to 
systems composed of multiple rods. If the rod spacing- 
to-length ratio is large (spacing > > length), then (50) 



R = 



p 22 

(In-), 



n27if 



(7.8) 



where n is the number of rods and the other variables are 
as previously defined. If the rod spacing-to-length ratio is 
small (length > > spacing) (50), 



R = 2^ (ln A ) ' 



(7.9) 



where A = (a S 12 S 23 S 34 



) l 



and S 12 = spacing between electrodes 1 and 2, and so 
forth. 





Resistivity, 


Resistance, 












XI 










rlOO 


r 




p5,000 




100 






-80 

1 co 100,000 








r3,000 


<u 








-60 








-1,000 


o 

c 


50 


- ^V 




-40 






- 1,000 


- 500 


D 
0) 


40 


VV Actual resistance 










- 500 




to 
<u 


30 


N \>^ with mutual effects 


Rod 

length, 


10,000 
^20 








- 100 


1 


20 


\ \ 

\ \ 


ft 


^s. 






r 100 


- 50 

. ft 


<u 








- io W»o 


- 




^.50 " 


Rod ratio, jpj 


c 


10 


S \ 




- 8 








- 10 


a* 




Expected ^S \ N. 




— 6 

100 

- 4 

L 30 






- 10 


uf 
o 

-z. 

CO 


5 
4 

3 


resistance \ N. 


Tur 


ling 


Ld 

rr 


2 


S 
\ 

V 


point 




1 


1 1 1 1 1 , . , 1 N J 


Figure 7.20.— Nomogram to provide resistance of driven rod. 


2 3 4 5 10 50 


Rod ra 


tio is equal t< 


) ro 


d length 


i (feet) dividec 


1 by rod diameter 






NUMBER OF RODS 



(inches). Example is shown for a rod of %-in diameter and 20ft 
length, driven in 500-Q-ft soil resistivity, providing about 35-Q 
resistance. 



Figure 7.22.— Resistance of parallel rods when arranged in 
straight line or circle with spacing equal to rod length. 



ROD LENGTH, ft _+ 20 18 




200 400 600 

RESISTANCE, XL 



15 




14 




13 




12 




11 




10 






c 


y 


o 






8 




7 


^_ T3 




o t= 


6 


5? 2 




en 


b 


LU o> 




o p 




2 i 




< "> 




1- o 




CO " 




CO 




LU 




en 



800 



Figure 7.21.— Resistance of one ground rod, %-in diameter. 





A, 0.5 ft B C 


20 


Jyy 


25 


/// /^ 




/// s' £",100 ft 


30 


'Wy/' ^^r^^' 


40 


JS^^^^ 


50 

70 


^y/ — A, B, C, D, E are various 

// rod spacings between 
/^ 0.5 and 100 ft 


1 00 


/ i i ii.i 



2 4 6 8 

NUMBER OF GROUND RODS 



10 



Figure 7.23.— Variation of earth resistance as number of 
ground rods is increased for various spacings between rods. 



169 



A formula for determining the resistance of grounding 
meshes is given by (53) 



""SKlS- 



+ K 



<B)° 



- k 2 ), (7.10) 



where L = total length of buried conductor, 

z = burial depth, 
and B = area enclosed inside mesh perimeter. 

The constants k x and k 2 depend upon the burial depth and 
the length-to-width ratio of the mesh and may be deter- 
mined from the graphs shown in figures 7.24 and 7.25 (53). 
However, for a typical mesh where the length and width 
are similar and the burial depth is a few feet or less, then 
k x = 1.3 and k 2 = 6. 

In many cases, combinations of rods and a mesh are 
used, especially when driven rods are interconnected by 
bare conductors that are also buried in the soil. For these 
situations (53), 



R = 



R X R 2 - R m 
Rj + R 2 — 2R n 



(7.11) 



where R x = rodbed resistance, 
and R 2 = mesh resistance. 



The mutual resistance, R m , is (53) 



R m = ± (In y + k, ^p - k 2 + 1), (7.12) 



where L = total length of mesh conductor, 
and i = length of one rod. 

If the soil is of uniform resistivity, adding a mesh to a 
preexisting rodbed, or vice versa, cannot be justified 
merely from the viewpoint of reduced resistance, since the 
reduction in resistance will seldom amount to more than 
10% to 15%. However, the addition of a mesh to a rodbed 
will usually smooth out the potential-gradient distribu- 
tion, and the addition of a rodbed to a horizontal mesh 
generally attenuates seasonal fluctuations in resistance 
(23, 55). 

Other electrode configurations are in use but are not 
as widespread as the two covered above. Table 7.3 summa- 
rizes most of the other electrode types and gives formulas 
for determining their resistance (22). As a first approxi- 
mation, the Laurent formula gives a quick and fairly 
accurate estimate of the ground resistance of any type of 
system (56): 



R -c + 



(7.13) 





1.4 










JC 


1.3 










h-' 










.4 


UJ 


1.2 










u 












u. 

Ll 


1.1 






B 




III 




. 








( > 












u 


1.0 
.9 


I 


C 


\ i 


i,i, 



KEY 
Curve A ■ For depth z = 

Curve B : For depth z = iSEgU 



. (area) l/2 



Curve C '■ For depth z = 



13 5 7 

LENGTH-TO-WIDTH RATIO 

Figure 7.24.— Values of coefficient k, as function of length- 
to-width ratio of area. 



UJ 

y 
u. 

UJ 

O 

o 



KEY 

Curve A : For depth z = 

^ o . i- j xl. (area) l/2 

Curve a . For depth z - — ^ — 



Curve C ■ For depth z 



_ (area)' 72 



13 5 7 

LENGTH-TO-WIDTH RATIO 



Figure 7.25.— Values of coefficient k 2 as function of length- 
to-width ratio of area. 



where L = total length of buried bare conductor, 
and r = equivalent radius of the system. 

The equivalent radius of a grounding system varies de- 
pending upon the exact configuration, but a safe estimate 
is one-half of the length of the longest diagonal line 
contained by the system (55). 

Contact resistance between the surface of the elec- 
trodes and the soil is not normally a significant factor if 
the bed has been in existence long enough for the soil to 
settle and compact, but in new beds it may amount to 20% 
of the total resistance (57). 

In summary, the best way to achieve a low-resistance 
ground is to maximize the periphery or areal extent of the 
grounding system. Conductor diameter has little effect 
upon resistance, and mechanical strength requirements 
should be the primary consideration. Because of wide 
seasonal variations in the soil resistivity of surface layers, 
deeply buried meshes or deeply driven rods are often 
preferable. This is also advisable if lower resistivity layers 
are known to exist at depth. Driven rods are usually 
preferred over buried meshes for three reasons (39): 

• The expense of earth removal to bury the mesh is 
avoided. 

• Rods do not require the packing of earth around the 
buried electrodes to ensure good earth contact. 

• The use of rods can give a desired resistance more 
easily than using any other ground-bed form. 

Note that although formulas are excellent for calculating 
the theoretical resistance of a grounding bed, the actual 
resistance should always be measured with an earth tester 
to ensure system integrity. 



170 



Table 7.3.— Approximate resistance formulas for various electrode configurations 



Electrode 
configuration 



Description 



One ground rod; length L, 
radius a 



*-£:»"*-« 



Two ground rods; spacing 
s>L 



= 7 £ r(ln «:_ 1)+ _£_ (1 .j4 + l4. . ., 



4TTL 



41TS 



1 2 c 4 

3s 5s 



• • 



Two ground rods; spacing 
s<L 



. ^£_ (ln Hk + ln *£ _ 2 + JL _ 



4TTL 



2L 16L 2 512L 4 



Buried horizontal wire; 
length 2L, depth s/2 



R = _£_ (ln ik +ln *k_ 2+ JL 



4TTL 



2 4 

a <5 

+ 



2L 16L 2 512L* 



A 
+ 

o 



Right-angle turn of wire; 
length of arm L, depth s/2 



Three-point star; length of 
arm L, depth s/2 



Four-point star; length of 
arm L, depth s/2 



Six-point star; length of 
arm L, depth s/2 



Eight-point star; length of 
arm L, depth s/2 



Ring of wire; ring diameter 
D, wire diameter d, depth s/2 



R = -£- (ln — + ln — - 0.2373 + 0.2146 £■ + 0.1035 Ar - 0.0424 2-r 
4TTL as L L 



R = -Sr (In — + In — + 1.071 - 0.209 f + 0.238 ^r - 0.054 S-... . .) 
6ttL as L ,2 T 4 



2 4 

r (In 2k + ln 2k + 2.912 - 1.071 £ + 0.645 ^ " °- 145 ^ • ■ ■) 



8ttL v a 



2 4 

R = rrg— (ln — + ln — + 6.851 - 3.128 f + 1.758 4* - 0.490 % 
12TTL as L .2 4 

2 4 

R = -rk-r (ln — + ln — + 10.98 - 5.51 f + 3.26 4t - 1.17 ■%■ . . 
16TTL a s L ,2 T 4 



d P /■, 8D 4D, 

R = — "=— (ln — + ln — ) 

2W 2 D d 



Buried horizontal strip; 
length 2L, section a by b, 
depth s/2, b, a/8 



D £_ ,, 4L 

R _ 4WT (ln "" + 



ab . , 4L . 
r + ln 1 



2(a + b) 



Buried horizontal round 
plate; radius a, depth s/2 



2 4 

r. - P I P n 7a I 33a •, 

R " 8a + 4tts (1 19 2 + /n 4 • • • ) 

12s 40s 



CZ^ 



Buried vertical round 
plate; radius a, depth s/2 



-^. + 7 £_ (1+ J5L+- 22 4. • •) 



8a 4tts 



2 4 

24s 320s 



Two-Layer Earth Structures 

In many situations, the soil is not homogeneous but 
consists of two or more distinct layers that are approxi- 
mately horizontal and possess differing resistivity values. 
The effect of a two-layer structure upon ground resistance 
depends upon the top-layer thickness, the relative conduc- 
tivity of the two layers, and the dimensions of the ground- 
ing system with respect to the thickness of the first layer 
(26). Figure 7.26 shows the potentials and potential gradi- 
ents for a mesh system in the first layer of a two-layer 
configuration where the thickness of the first layer ranges 
from 0.1 to 1,000 m (18). In this case, the first-layer 
resistivity (pi) is 200 Q-m, and p2 is 600 fi-m. The equiva- 
lent radius of the grounding grid is 10 m, and the reflec- 
tion factor (K), as defined by the following equation, is 0.5: 



K = 



p2 - pi 
p2 + pi - 



(7.14) 



It may be seen that the potential gradient depends almost 
solely upon the first-layer resistivity if the grounding 
system is wholly immersed in that layer. The effect of 
first-layer resistivity upon ground-bed resistance in- 
creased with the thickness of that layer. Thus, if pi <p2, 
ground resistance will decrease as top-layer thickness 
increases. 

Soil-Heating Effects 

The manner in which a ground bed responds to the 
flow of current through it depends upon the magnitude 
and duration of the loading. Two types of loading have 
been recognized and will be dealt with separately. 

Long-term loading of the safety ground bed in a mine 
power system should consist only of currents due to 
unbalance, the charging of conductor capacitances, and 
mutual inductance between conductors. At any rate, in a 
properly functioning system, the current magnitude 



171 



Mesh 
design 




KEY 

Curve A ■ Potential rise of mesh electrode above 
remote ground 

Curve B '■ Potential rise of center of mesh above 
remote ground 

Curve C '■ Potential difference between center of 
mesh and mesh electrode 



0.1 I 10 100 1,000 

FIRST-LAYER HEIGHT, m 

Figure 7.26.— Influence of first-layer height of potentials. 



should be on the order of a few amperes. If the bed is very 
extensive, the dissipation of ground current in the soil may 
cause only a small rise in soil temperature. Because of the 
negative temperature coefficient of soil, the actual ground- 
bed resistance will decrease (23). If the temperature rise is 
high enough to evaporate some soil moisture, then the 
resistance will increase somewhat. Capillary action will 
tend to restore any moisture, and the soil itself will also 
conduct away some of the heat. Eventually an equilibrium 
point will be reached where the system is once again 
stable, although the soil temperature and ground-bed 
resistance may be slightly altered. The maximum allow- 
able ground-bed current is given by (50) 



low rate. In this situation, the maximum allowable soil 
temperature rise is given by (57) 



I = ^ (2pX0) 1/2 , 



(7.15) 



where p = soil resistivity, Q-m, 

X = soil thermal conductivity, 1.2 W/(m- °C), 
and 6 = maximum allowable soil temperature rise, 
°C. 

If both sides of the equation are multiplied by R, the 
maximum permissible applied voltage is found to be 



V = (2p\ef 



(7.16) 



Generally, the maximum allowable temperature is 100° C, 
at which point total evaporation occurs. Therefore 6 may 
be replaced by (100 - T) where T is the ambient Celsius 
temperature. The preceding analysis is subject to two 
restrictions (57): 

1. The thermal conductivity, X, is somewhat tempera- 
ture dependent, and 

2. Soil moisture will start to evaporate at tempera- 
tures below 100° C. 

Short-term overloading of the grounding system may 
occur during certain fault situations, but in a properly 
functioning system, only the grounding conductors, lo- 
cated inside cables and with the overhead powerlines, and 
the neutral resistor should be subjected to fault current. 
Should a situation occur in which the ground bed is called 
upon to handle large currents for a short time, heat 
conduction through the soil may be ignored because of its 



0.24i 2 pT 
da 



(7.17) 



l 

T 
S 



where p = soil resistivity, Q-m, 

= current density at electrode surface, A/m 2 , 

= time, h, 

= soil density, kg/m 3 , 

and a = soil specific heat, kWh/°C - kg. 

So far only the effects of ac upon soil heating have 
been discussed. Dc causes completely different phenom- 
ena. The first of these is polarization. The flow of dc 
through water causes some of the molecules to dissociate 
into the constituent gases, hydrogen, and oxygen. The 
resulting gas bubbles eventually form a film on the 
electrode surfaces, thereby insulating them from the sur- 
rounding soil, which leads to a dramatic rise in resistance. 

In addition, dc causes electro-osmosis (also referred to 
as endosmosis). Here, moisture present in the soil (which 
is not electrolyzed) tends to migrate toward the negative 
electrode of the dc source. Actually, cations present in the 
soil are attracted to the cathode, and the polar water 
molecule is normally attached to these positive ions. 
Again, an increase in resistance is the result. 

Control of Potential Gradients 

In addition to providing a low-resistance path to 
ground, the ground bed should also be designed so that 
potential gradients in the soil surrounding the bed (step 
and touch potentials) are held to a minimum for the 
protection of personnel. 

As a generalization, it can be stated that meshes are 
superior to rodbeds as far as potential-gradient control is 
concerned (18, 23). This is illustrated by table 7.4, which 
compares a variety of grounding systems, each having 
about the same total length of buried conductor (18). The 
electrodes are buried to a depth of 0.6 m, and as can be 
seen, grid C (rodbed) shows significantly higher potential 
than does grid A (mesh). The potential gradients around a 
mesh may be decreased by making the meshes smaller. 
Figures 7.27 and 7.28 show the improvement which can be 
obtained by burying the grounding system to a greater 



172 



Table 7.4.— Comparison of grounding grids with other types 
of electrodes 



Grid 



Maximum 

Rod or Total len 9 th Lenath mesh 

_ . of buried , ~. Number voltage, 

mesh . , of rods, . . „. , 

. , conductor, _ of rods % of 

avnnt ' m 



potential 



V 




/■ 




X 




/ 




- 



• • • • • 

• • • • • 



160 



174 



175 



NAp 



NAp 



25 



18.9 



12.7 



32.4 



100 


A B C D E F 












80 




f ' 








o-° 








± 






VOLTAGE, 
o 


T o (J 
\ l2 




i 

j 


-e— Rod 


y 40 

a. 
a. 
< 

20 


-\a 




2.4 U 

3.6 U 
4.8 

6.0 Burial depth, 
from surface to 
s. top of rod, ft 




/ E 


■ 





, i , i 







5 10 


15 




DISTANCE 


FROM ROD, ft 





Figure 7.27.— Potential on ground surface due to rod 6 ft 
long and 1-in diameter buried vertically at various depths. 



176 



176 NAp 



16 23.2 



17.6 



depth (24). It is obvious that the deeper a bed can be 
buried, the better will be the gradient control. A rodbed, 
where the rods are interconnected by bare conductors with 
the entire system buried to a depth of a few feet, should 
provide both a low resistance and a safe potential gradi- 
ent. Building a fence around the perimeter of the ground 
bed is one way of limiting human exposure to hazardous 
potential gradients. 



GROUND-BED RESISTANCE MEASUREMENT 



KEY 
Length of electrode: A 30 ft 
5 20ft 
C 10 ft 
D 5ft 



of strip 




8 12 

DISTANCE FROM 0, ft 

Figure 7.28.— Potential on ground surface due to strips, 1 in 
by 0.1 in, of various lengths buried horizontally at depth of 2 ft; 
values given are those along line OY perpendicular to length of 
strip. 



Measurement Method 

The accepted technique for determining the resis- 
tance to infinite earth of a grounding resistance is called 
the fall-of-potential method (36). Figure 7.29A shows a 
drawing of this arrangement (56). Three terminals are 
required: the ground under examination, a potential elec- 
trode, and a current electrode. The current electrode is 
spaced far from the ground system being tested, and the 
potential electrode is placed at some point on a straight 
line between the two. The resistance-measuring equip- 
ment is operated, and a reading is taken. Here, a known 
current is passed through the current electrode, the volt- 
age between the potential electrode and ground is mea- 
sured and the resistance is the ratio V/I. This process is 
repeated as the potential electrode is moved farther and 
farther from the grounding electrode, toward the current 



electrode. A graph is then drawn in which the ground 
resistance is the ordinate and the distance between the 
ground and potential electrodes is the abscissa. Figure 
7.29B shows two typical plots that may result (56). Curve 
a was taken with the current electrode at a greater 
distance than in curve b. The flat portion of curve a is an 
indication that the current electrode is now far enough 
away from the grounding system that the mutual effect no 
longer exists. This is illustrated in figure 7.30 by the 
hemispheres of influence surrounding the ground and 
current electrodes (35). 

The proper spacing for the measurement probes is 
based upon hemispherical electrodes, so any actual ground 
system must first be converted to an equivalent hemi- 
sphere before the needed spacing can be determined (56). 
This may be approximated by assuming that the equiva- 
lent radius is equal to one-half the length of the longest 



173 



Current source 



Ammeter 



O^D 



y 



Voltmeter 



Ground 



Potential 
electrode 



1 Current 
electrode 



i 

h c 



A Fall-of-potential method 




DISTANCE (P) FROM GROUND TO POTENTIAL 
ELECTRODE (PE) 

B Earth resistance curves 

Figure 7.29.— Measuring resistance of grounding system. 



Megohmmeter 




Potential 
electrode 



Current 
electrode 



Grounding 
electrode 



Figure 7.30.— Concentric earth shells around ground con- 
nection being tested and around current electrode. 



diagonal that can be placed inside the perimeter of the 
system (that is, 50% of the maximum bed dimension). 
Figure 7.31 shows the proper spacing for both current and 
potential electrodes for a given equivalent radius of the 
grounding system (55). The potential-electrode spacing 
that yields the true value of ground resistance is equal to 
about 61.8% of the current-electrode spacing. For large 
ground systems, it may be impossible to attain the neces- 
sary spacings for potential and current electrodes result- 
ing from this technique. In that case, the procedure 



200 r- 



160 - 



120 - 



80 



40 



L±J 
O 



Distance to 
current 
electrode 




Distance to 
potential 
electrode 



10 



£ 800 r 



600 



400 - 



200 



Distance to 
current -v. 
electrode ^ 




Distance to 
potential 
electrode 



_L 



20 40 60 80 

RADIUS OF EQUIVALENT HEMISPHERE, ft 



100 



Figure 7.31.— Correct spacing of auxiliary electrodes to give 
true resistance within 2.0%. 



outlined earlier may still be followed, that is, varying the 
potential electrode spacing while keeping the current 
electrode at some fixed spacing as far as possible from the 
grounding system. The true resistance may then be de- 
rived from the resulting graph using one of several avail- 
able methods (56*). 

Ground Test Instruments 

Certain precautions should be observed when a 
ground test instrument is chosen. A machine that uses dc 
should be avoided because of problems with polarization 
and electro-osmosis. Ac is satisfactory, but a frequency 
slightly removed from the actual power frequency is pref- 
erable so the effects of stray currents can be avoided. On 
the other hand, if the frequency used is too far removed 
from the power frequency, erroneous results may occur 
since ground resistance (impedance) varies with frequency 
(45). The leads from the instrument to the electrodes 
should be spaced as far apart as possible to minimize the 
effects of mutual inductance and capacitance. In a good 



174 



instrument, the resistance of current and potential probes 
is not critical, but inferior equipment will give readings 
that vary widely depending upon the probe resistance. 
Great accuracy in measuring earth-ground resistance is 
not critical because the earth resistance measurement 
techniques themselves can never be precise or accurate. 



Table 7.5.— General resistivity classification 

Conductivity characteristic of material Resistivity, Sl-cm 

Good 10~ 3 —10 

Intermediate 10 2 —10 9 

Poor 10 10 —10 17 



GROUND-BED RESISTIVITY 

In the discussions on resistance it was pointed out 
that soil resistivity, p, is an important parameter; specifi- 
cally, ground-bed resistance is directly proportional to soil 
resistivity. The resisitivity of a material was defined in 
chapter 2 as the resistance in ohms between the opposite 
faces of a unit cube of that material. The value of resistiv- 
ity varies widely depending upon the substance being 
measured; for rocks and minerals, it may range from 10 ~ 3 
to 10 17 Q-cm. A general classification is shown in table 7.5 
(19). Efforts have been made to relate resistivity values to 
the geologic age of various rocks, as can be seen in table 
7.6. As a rule, resistivity increases with rock age (5), but 
there are exceptions (54). 

Rock structure enters into resistivity determinations, 
in addition to geologic age. The resistivity of a newly 
formed rock depends mainly upon the amount of water it 
contains. Young rock will generally have a large pore 
volume and hence a fairly significant quantity of connate 
water; therefore, it will exhibit a low resistivity. As time 
passes and the rock is subjected to forces that tend to 
consolidate, compress, or metamorphose it, the pore vol- 
ume and water content will decrease, with a subsequent 
increase in resistivity (5). Hard crystalline rocks are 
usually bad conductors, but if crushed or badly fractured, 
their resistivity may decrease because of greater porosity 
(47). Resistivity values for some common soils are given in 
table 7.7 (55). 

When completely dry, most rocks and minerals are 
nonconducting, although some metallic ore bodies will 
carry current (24). The main soil constituents have very 
high resistivities, and in fact, the oxides of silicon and 
aluminum are good insulators (50). Figure 7.32 reviews 
the resistivities of some common rocks, ores, and metals 
(47). 

Factors Affecting Resistivity 

Several factors can affect resistivity, and these are 
generally considered to include 

• Moisture content, 

• Dissolved salts, 

• Temperature, 

• Soil type, 

• Grain size and distribution, and 

• Location. 

The level of influence for each is described in the following 
paragraphs. 

Soil containing no moisture has a very high resistiv- 
ity. The addition of water causes a sharp increase in 
conductivity, but the decrease in resistivity rapidly levels 
off once the moisture content of the soil reaches about 16 
wt %, as shown in figure 7.33 (55). Tests by the Bureau of 
Standards have indicated that resistivity increases mark- 
edly when moisture content falls below 20% (36). 



Table 7.6.— Variations in resistivity with geologic age 



Approxi- 
mate 

resistivity, 
fi-m 


Quarter- 
nary 


Creta- 
ceous, 
Tertiary, 
Quarter- 


Pennsyl- 
vanian, 
Missis- 
sippian, 


Cambrian, 
Ordovician, 
Devonian 


Pre- 
Cambrian 

and 
combina- 
tion with 






nary 


Triassic 




Cambrian 


1 

5 




Loam 








10 




Clay 








20 




Chalk 


Chalk 






30 






Trap 






50 






Diabase 






100 






Shale 






300 






Limestone 


Shale 




500 






Sandstone 


Limestone 




700 










Sandstone 


1,000 


\ Coarse 






Sandstone 




1,200 


I sand 








Quartzite 


1,500 


land 






Dolomite 




3,000 


/gravel in 








Slate 


5,000 


I surface 








Granite 


10,000 


/ layers 








Gneiss 



Table 7.7.— Typical values of resistivity of some soils 

Type of soil Resistivity, Q-cm 

Loams, garden soils, etc 500- 5,000 

Clays 800- 5,000 

Clay, sand and gravel mixtures 4,000- 25,000 

Sand and gravel 6,000- 10,000 

Slate, shale, sandstone, etc 1,000- 50,000 

Crystalline rocks 20,000-1,000,000 



Copper 

Silver 

Gold 



T 1 1 1 1 r- 

Wet limestone E2 

Chromire ore I 



Rock salt 



Marble 

quartz 

V////////K 

. Zinc-blend 
ore 



Y////////A Wet-to-moist 

granite, granuhte 

Moraine Y////////A 



Iron 

I Zinc 

Lead 



UZZZ2 Clays 

v//;;;/;;;;;/» Hematite ore 

V//////////////A Galena ore 
V///////////A Magnetite ore 
Y////////////////K Pyrite ore 
^ Graphite Y/////////A Graphitic shales 

£Z) Psilomelane, hollandite, pyrolusire 
B2 Pure chalcopyrite 
E3 Pyrrhotite 



J_ 



10 



8 



10" 



I I0 4 

RESISTIVITY, Sl-n\ 



I0 8 



10'' 



Figure 7.32.— Resistivity range of some rocks, minerals, and 
metals. 



175 



200 r 




8 16 24 

MOISTURE CONTENT, wt % of dry soil 

Figure 7.33.— Variation in soil resistivity with moisture con- 
tent. 



The conductivity of water is not a constant value, and 
it has noticeable effects on soil resistivity. Very pure water, 
such as may be found high in the mountains, has a poor 
conductivity, and as a result, mountain soil may be very 
wet and still possess a high resistivity (24). To a large 
extent, it is the dissolved salts present in the water that 
make the solution conductive. Conduction is electrolytic in 
nature; that is, current flows via the movement of positive 
and negative ions in solution. Thus, the concentration of 
dissolved salt, the particular type of salt, and the solution 
temperature all have an influence upon the degree to 
which a dissolved salt can lower soil resistivity. Figure 
7.34 shows the effect of various salts upon resistivity (55). 

Water has a large negative temperature coefficient of 
resistivity, and the transition from liquid to solid state is 
marked by a dramatic rise in resistivity (31). In addition, 
most electrolytes have a negative temperature coefficient 
of resistivity, amounting to about - 2.0%/°C (24). Table 7.8 
illustrates this effect (34). 

Table 7.8.— Effect of temperature on resistivity of water 



Temperature, 

OQ1 

20 

10 

(liquid) 



Resistivity, 
Q-cm 

7,200 

9,900 

13,800 



Temperature, 
°C' 



(ice) . 

-5 

-15... 



Resistivity, 
Q-cm 

30,000 

79,000 

330,000 



1 To convert to degrees Fahrenheit, multiply by 9/5 and add 32. 

When a very high impulse current such as a lightning 
stroke enters a ground bed, the resulting voltage gradient 
may be so high that the soil breaks down. These current 
levels can be extremely damaging to the soil. Lower 
current levels flowing into a ground system for extended 



10,000 r 



5,000 



o 

o 

o 



KEY 

Copper sulfate 
Sodium sulfate 
Sodium carbonate 
Sodium chloride 
Calcium carbonate 
Sodium hydrate 
Sulfuric acid 



> 

h- 

00 

co 

LU 

cr 



1,000 



500 



100 




0.20 



0.08 0.12 

SOLUTION, % 

Figure 7.34.— Typical resistivity curves of solutions. 



periods may heat the soil to the point where most of its 
moisture will evaporate. When this condition is reached, 
soil resistivity increases drastically. 

Different soils are characterized by various resistivity 
levels (table 7.7). To a large extent, this is due to the 
previously discussed effects of structure as it pertains to 
conductivity. Loams and clays possess a low resistivity, 
while shales, sandstones, and crystalline rocks occupy the 
high end of the scale (50). 

The nature of the particles making up the soil or rock 
is another aspect of rock structure, which influences 
conductivity through the rock's ability to trap and retain 
water. Surface tensions cause water to cling to large soil 
particles or grains; with small-grained substances, mois- 
ture simply fills up the multitude of pore spaces between 
individual particles. The range of particle sizes and their 
packing determines how much of the volume occupied by a 
particular soil will be void space and thus available for 
filling by water. If most of the grains are the same size, 
total pore volume may range from 26% to 46%, depending 
upon the manner in which the grains are packed (19). 

If a particular rock structure or formation is confined 
to a small geographical area, then it probably has a fairly 
uniform resistivity, excluding areas of subsequent igneous 
activity. Should the formation be widespread, however, 
chances are that variable resistivities will be noted de- 
pending upon location. This is due to the differences in 
local conditions that may have prevailed over a small area 
during actual deposition or formation of the rock strata. 
This may also be caused by variations in the ground water 
properties from place to place within a large region (5). 

Resistivity Measurements 

The basic procedure for measuring soil resistivity 
involves the determination of the potential gradient on the 
earth's surface caused by flow of a known current through 
the area. 

To illustrate the basic technique, assume an earth 
structure composed of two horizontal layers, the top one of 



176 



high resistivity, p lt and the lower one of low resistivity, p 2 , 
as shown by figure 7.35 (19). The thickness of the upper 
layer is given by h. A power source forces current flow 
through the ground between the two outer electrodes. At 
very small electrode spacings, the apparent resistivity will 
approximate p 1 since most current flow would be confined 
to the upper layer. At very wide spacings (much larger 
than h), the apparent resistivity will be about the same as 
p 2 , because the majority of the current would flow through 
the deeper layer. 

Many methods are available for measuring earth 
resistivity, such as the techniques of Gish-Rooney, Lee, and 
Schlumberger. Most of these procedures are based on the 
arrangement described by Wenner (58), which is shown in 
figure 7.36 (35). Four uniformly spaced electrodes are used, 
and a current source is connected across the two outer 
terminals while the potential drop is measured across the 
inner terminals. When the electrode length b is small 
compared with the spacing a, then the resistivity is (51) 

p = 2iraR, (7.18) 

where p = resistivity, fl-m or Q-ft, 

a = spacing between electrodes, m or ft, 
and R = resistance = V/I, fl. 

Some problems that may arise from the use of this method 
are 

• Stray currents due to leakage as from motors, 

• Natural currents due to electrolysis of nearby min- 
erals, 

• Polarization due to use of a dc source, 

• Inductance between the lead conductors, and 

• Leakage from the conductors and the instrument 
when in wet areas. 

The first three problems are circumvented through the use 
of an ac source operating at the nonpower frequency of an 
instrument that generates the equivalent of a square 
wave. The use of a well-insulated instrument and conduc- 
tors solves the latter two difficulties. The megohmmeter 
has all these features and is an excellent apparatus for use 
in work of this type. 

To perform a resistivity survey, the megohmmeter is 
set up as shown in figure 7.36, the instrument is operated, 
and a resistance value R is read from the built-in meter. 
The procedure is then repeated at different electrode 
spacings. A graph may be made comparing the resistivity, 
p, with the electrode spacing, a, as shown in figure 7.37 
(55). For each value of electrode spacing, there is a corre- 
sponding value of resistivity, p a , seen by the instrument. 
This apparent resistivity is equal to the resistivity that a 
semi-infinite homogeneous earth would display at an 
equal electrode spacing and an identical value of R. In the 
example shown, the apparent resistivity decreases as 
electrode spacing increases. The overall shape of the curve 
indicates that the soil here is composed of two horizontal 
layers, with the overlying horizon having a higher resis- 
tivity then the lower one. As the electrode spacing, a, is 
increased, more and more of the current flow between the 
outer electrodes occurs in the deeper layer of the soil, and 
this is reflected in the continuous decrease in the apparent 
resistivity (5). 

In a case like the one just described, a grounding grid 
composed of deeply driven vertical rods would be best, 
since the rods would penetrate into the underlying layer of 




Electrodes 



', >P 2 



Figure 7.35.— Diagram for four-electrode resistivity survey 
showing lines of current flow in two-layer earth. 




Megohmmeter 









Figure 7.36.— Connections for Wenner four-terminal 
resistivity test using megohmmeter; distance a should be at 
least 20 times b. 



in 

UJ 

or 



UJ 

or 
< 
a. 
a. 
< 




ELECTRODE SEPARATION 

Figure 7.37.— Typical curve of resistivity versus electrode 
separation. 



177 



higher conductivity and thus provide a more effective 
ground. Additionally, soil horizons near the surface are 
usually subject to wide seasonal variations in resistivity 
due to changes in ambient temperature and moisture (40). 
Tagg (55) presents several methods whereby an accu- 
rate interface-depth determination may be calculated. 
Values are read from a standard graph, and multiple 
calculations are then performed, followed by another 
graph construction from which the correct depth is read. 
Core drilling has verified that values derived in this 
manner agree closely with the actual conditions. 

Effect of Chemical Treatment of Soils 

The natural resistivity of some soils is so high that it 
is virtually impossible to construct a ground bed with a 
satisfactorily low value of resistance. By injecting into the 
earth a substance whose resistivity is very low, the local 
soil resistivity can be effectively reduced, thereby lowering 
the resistance of a grounding grid. Such chemical treat- 
ment acts to increase the apparent dimensions of the 
metallic electrodes (7). The result of chemical treatment is 
to reduce ground resistance by a considerable amount, 
often as much as 15% to 90%. Figure 7.38 shows an 
example of this effect (36). Generally, the percentage 
improvement is greater for a very high resistance ground. 

Substances traditionally used as chemical additives 
include sodium chloride, calcium chloride, copper sulfate, 
and magnesium sulfate (36). Newer additives include gels 
composed of acrylamide, silicic acid, or copper ferrocya- 
nide. In the past, electrodes were sometimes surrounded 
by a bed of coke, not a true chemical treatment but rather 
a partial soil substitute (24). The effectiveness of most 
treatments in lowering ground-bed resistance is about the 
same, with the ultimate selection depending upon the 
criteria of cost, availability, and corrosive properties. 

A prime disadvantage shared by most chemical treat- 
ments is the fact that they will corrode most metals (7). 
Magnesium sulfate has little or no corrosive effect, and 
graphite is also innocuous. Other additives generally 
speed up the decay of grounding electrodes. 

Another disadvantage is that chemical treatments 
are dissipated and carried away by neutral drainage 
through the soil (36). Acrylamide gel, which is not water 
soluble, is an exception (34). The rate at which chemical 
additives are washed away depends upon the soil type and 
porosity as well as the amount of rainfall. Useful life may 
range from 6 months to 5 or more years. 

The cost of chemical treatment may be higher than 
the price of driving longer ground rods to reach deeper, 
lower resistivity soil layers, but in some instances it is not 
feasible or desirable to increase penetration depth. As 
shown in figure 7.39, the seasonal variations in resistance 
that are exhibited by grounding grids because of temper- 
ature and moisture fluctuations, are attenuated in those 
cases where chemical treatment has been applied (36). 

The best method of application, illustrated in figure 
7.40, is to dig a circular trench about 1 ft deep and with an 
inside diameter of 18 in around each ground rod (36). The 
additive is placed into the trench and then covered with 
earth. The area is then flooded with water to initiate the 
solution process. In this manner, the solution can perme- 
ate a greater volume of soil, while any corrosive action is 
minimized. 



1,600 r # 



Before treatment 




July 



July Jan. 

MONTH 



July 



Jan. 



Figure 7.38.— Reduction in ground mat resistance by soil 
treatment. 




Untreated 



Treated 

j i I L 

Jan. 
MONTH 

Figure 7.39.— Seasonal resistance variations attenuated by 
soil treatment. 




Ground 
rod 



Figure 7.40.— Trench model of soil treatment. 



GROUND-BED CORROSION 

Corrosion is a phenomenon that must be considered in 
the design of a ground bed. There are three basic ways by 
which underground corrosion can occur (52): 

• Dissimilar metals connected together electrically 
and surrounded by an electrolyte such as soil, 



178 



• Dissimilar electrolytes in close proximity to the 
same piece of buried metal, and 

• Stray electrical current leaving a buried metal 
structure. 

In the first mode, variations in electrochemical poten- 
tial provide the key to the dilemma. The standard half- 
cell, upon which most corrosion work is based, consists of 
a copper rod bathed in a saturated copper sulfate solution. 
When measured with reference to the copper and copper 
sulfate half-cell, each metal displays a certain character- 
istic potential, as shown in table 7.9 (61). If two metals are 
joined and immersed in soil, the one whose potential is 
more negative will discharge current and be corroded, but 
the more positive (noble) species will collect current and be 
protected. When only one metal is used, corrosion can still 
occur because of differences in soil composition. Metal in 
an oxygen-rich zone will be protected, while metal in a 
relatively oxygen-poor soil horizon will be attacked. For- 
eign metallic structures in the grounding-grid vicinity, 
such as pipes, cable sheaths, and building frames, may 
also act in conjunction with the ground bed to form an 
anode-cathode corrosion situation. 



Table 7.9.— Typical potentials of metals in soil measured from 
a copper and copper sulfate reference electrode 

Metal Potential, V 

Magnesium -2.5 

Aluminum -1.3 

Zinc -1.1 

Iron -.7 

Copper - .2 



The engineer designing a ground-bed system is faced 
with the problem of solving two conflicting sets of de- 
mands. For safe grounding, a very low resistance is desired 
between the soil and the buried metallic grid. To eliminate 
potential-gradient hazards, all metal structures should be 
tied together. However, protection requires that under- 
ground metallic structures be insulated from the corrosive 
effects of the soil. Similarly, the soil and metallic structure 
should be isolated from one another (61). This seeming 
paradox may be remedied by making the correct choice of 
ground-bed conductor and by applying suitable preventive 
techniques. 

Copper makes an ideal ground-bed conductor since it 
is corrosion resistant, has a high electrical conductivity, 
and is easy to clamp or weld (61). However, a good 
all-copper system is often ruined by tying it together with 
noncopper structures in the same locale, thereby leading 
to the corrosion of the less noble species (52). If the ground 
bed must be located in an area where steel or lead are 
present, two options are available. First, an insulating 
coating may be applied to the base metals. If this is not 
feasible, an all-steel grounding system is preferable, or one 
composed of steel rods connected with insulated copper 
wire (8). The idea here is to minimize the exposed surface 
area of the more noble metal. Normally, steel electrodes 
can be improved by applying a heavy zinc coating or by 
driving zinc electrodes in addition to the steel. Known as 
sacrificial anodes, the zinc conductors will be preferen- 
tially attached, thereby protecting the steel members. For 
extra protection, magnesium may be used instead of zinc. 
In highly corrosive soils, it may be necessary to utilize an 
external power source that supplies dc to the soil in order 
to nullify the natural corrosion currents. This is known as 



cathodic protection (41). For externally driven anodes, zinc 
or magnesium may be used; graphite and high-silica cast 
iron are also suitable. 

It may be seen that judicious choice of grounding 
materials and the use of corrosion-prevention techniques 
such as cathodic protection can provide a ground bed that 
is both low in resistance and high in longevity. 



GENERAL GROUND-BED GUIDELINES 

The primary objective of a grounding system is "to 
limit the potential rise above ground that appears on the 
frame and enclosures of the equipment connected to the 
power system" (30). Consequently, the station ground and 
safety ground beds should be spaced at least 50 ft apart, 
even though the law presently permits only a 25-ft sepa- 
ration (21). A typical voltage-gradient representation is 
shown in figure 7.41 (30). The two ground beds must be far 
enough apart so current surges through the station 
ground bed will not cause the safety ground bed to rise to 
more than 100 V above infinite earth. 

Once the site has been selected, the excellent guide 
developed by King (39) can be used for the design and 
construction of low-resistance driven-rod ground beds. The 
simplified procedure consists of the following four steps. 

1. Using the Wenner array, earth resistivity is mea- 
sured along the two lines at right angles to each other, 
centered across the proposed ground-bed site. Two mea- 
surements, with 6-ft and then 18-ft spacings, are taken 
along each line. 

2. Depending on the magnitude and homogeneity of 
the resistivities measured, the rod length, number, and 
arrangement are selected from tables. These tables (39) 
are based on the same information presented earlier in 
this chapter but are too extensive to be reproduced here. 



L±J 

o 
> 



100 V 




System 
ground 



Safety 
ground 



Figure 7.41.— Voltage gradients in earth during ground-fault 
conditions. 



179 



3. The selected rod configuration is driven, and the 
rods are interconnected with flexible, bare copper conduc- 
tors. Recommended size is 4/0 AWG, and connection 
should be clamped or brazed, but never soldered (1). 

4. The completed bed is measured by the fall-of- 
potential method to check that its resistance is below 5.0 
Q. If it is more, a new resistivity, p n , is calculated by 

P» = Po|. (7-19) 



where p = old resistivity, 

and R = measured resistance. 

Again, using the tables, additional rods are selected and 
then driven. Afterward, the resistance is again measured. 
Whatever the procedure used to construct the bed, the 
resistance should be checked not only when it is installed 
but periodically thereafter to ensure that it is still func- 
tioning properly. 



GROUNDING EQUIPMENT 

The basic resistance-grounded system consists of a 
resistance inserted between the power-system neutral 
point and ground. Specific concerns when selecting the 
grounding resistor are resistance, time rating, insulation, 
and connection. A problem also exists if there is no 
available connection to the power-system neutral. 

Grounding Resistor 

The ohmic value of the resistor is determined by the 
line-to-neutral system voltage and the maximum ground- 
current limit. As stated earlier, when portable or mobile 
equipment is involved, the maximum limit on low-voltage 
and medium-voltage systems is 25 A, and the upper 
current limit on high voltage is set by the grounding- 
conductor resistance, because flow through this conductor 
cannot cause any machine frame potential to be elevated 
more than 100 V above earth potential. However, high- 
voltage limits are typically chosen at 25 or 50 A (50 A is 
the maximum allowed in some States). For instance, if 
grounding-conductor resistance is 3.3 Q, and maximum 
allowable ground current is 30 A, then 25 A is normally 
chosen. When the resistance-grounded system is feeding 
only stationary equipment, there is no specified maximum 
ground current, but industry practice sometimes specifies 
low-resistance grounding with a 400-A limit. For all appli- 
cations, sizing the ohmic value of the grounding resistor is 
simply performed by dividing the line-to-neutral voltage 
by the selected ground-current limit; conductor impedance 
is neglected. The technique is justified by the method of 
symmetrical components for a line-to-neutral fault. 

Ground current can be limited at a level less than the 
restricted maximum, but for high-resistance grounding 
the smallest value chosen has two concerns: ground-fault 
relaying and charging current. For maximum safety, 
ground protective circuitry should sense ground current at 
a fraction of current limit (see chapter 9). Hence, reliable 
relay operation with electromechanical devices can be a 
problem if maximum current is less than 15 to 20 A. The 
other limitation is that ground-fault current should al- 
ways be greater than the system-charging current (59), the 
current required to charge system capacitance when the 



system is energized (see chapter 11). When very low 
ground-relay settings are used, the charging current may 
itself cause tripping. 

The second main concern in selecting a grounding 
resistor is its time rating, or the ability to dissipate heat. 
A grounding resistor carries only a very small current 
under normal system operation, but when a ground fault 
occurs, the current may approach full value. The high 
current exists until the circuit breaker removes power 
from the faulted circuit, which may take from a fraction of 
a second to several seconds depending upon the protective 
circuitry used. With correct fault removal, the physical 
size of the resistor can be small, as very little heat is 
produced. However, protection devices have been known to 
malfunction, and in these instances ground current might 
continue to flow until the power is removed manually. 
Thus, the resistor must be able to dissipate the power 
produced from full ground current for an extended time 
when portable or mobile equipment is involved. If not, the 
resistor can burn open and unground the system. Two 
ratings that ensure safety are continuous and extended 
time. These are essentially the same, since the extended- 
time rating refers to a heat-dissipation ability for 90 days 
per year (32). 

To provide a safety margin, the transformer-neutral 
side of the resistor (often called the hot side) must be 
insulated from ground at a level to withstand the line- 
to-line system voltage. Both resistor ends are at ground 
potential with normal operation but under a ground fault, 
the transformer end can approach line-to-neutral poten- 
tial. To afford good insulation, it is recommended that the 
resistor frame be placed on porcelain insulators, not tem- 
porary supports such as wooden blocks. Furthermore, for 
wye-connected secondaries, the transformer-neutral bush- 
ing must be insulated to at least line-to-neutral voltage. 

The last concern is the resistor connection. The 
grounding resistor is installed between the transformer 
neutral and the safety ground bed. In substations it is 
important to use insulated conductors, because bare con- 
ductors can easily compromise the required separation 
between the system and safety ground beds. Grounding 
conductors must extend from the ground-bed side of the 
resistor. Finally, to minimize resistor conductor lengths, 
the resistor must be located on the power-source end of 
distribution, as close as possible to the source power 
transformer. Distances greater than 100 ft are usually too 
long. 

Grounding Transformers 

Delta-wye, wye-delta, and delta-delta power trans- 
formers are extremely important in mine power distribu- 
tion because they offer very high impedance to zero- 
sequence currents. As a result, a ground fault existing on 
the secondary will do no more than raise primary line 
current. However if the transformer has a delta secondary, 
there is no neutral point to which the grounding system 
can be connected. Another case where this occurs involves 
mines where the utility company owns the substation and 
supplies ungrounded delta power. For both these situa- 
tions, a separate grounding transformer is needed to 
obtain an artificial neutral. The two types of grounding 
transformers in general use are the zig-zag and wye-delta, 
with the former being more popular. 

As shown in figure 7.42, the zig-zag is a special 
three-phase transformer designed for deriving the neutral. 



180 



The transformer winding interconnections are such that a 
very high impedance is shown for positive-sequence and 
negative-sequence currents but a very low impedance is 
exhibited during zero-sequence flow. 

A wye-delta grounding-transformer bank uses three 
identical single-phase transformers (fig. 7.43). The pri- 
mary windings, rated at line-to-neutral voltage, are con- 
nected in wye among the power-transformer secondary 
terminals and the grounding-resistor hot side, and the 
secondaries are connected in delta. Any secondary voltage 
rating can be used. Normally, no secondary current will 
flow, but during a ground fault, current will circulate in 
the secondary. This will cause the ground to be shared by 
the three transformers such that the neutral point will 
remain at constant potential. 

Grounding-transformer capacity only needs to be 
large enough to carry the maximum ground-fault current. 
Grounding transformers' primaries cannot be fused, as an 
open fuse will essentially unground the system, creating a 
dangerous situation. 



Main transformer 
delta secondary 



Zig-zag 
- transformer 



To ac circuit breakers 



Grounding 
resistor 
1 WW 



Figure 7.42.— Delta secondary with zig-zag grounding. 



Incoming power 



Wye-delta 
grounding 
transformer 



System 

ground 

bed 



Grounding 
resistor 




Safety -= 
ground 
bed 



Figure 7.43.— Delta secondary with wye-delta grounding 
transformer. 



SUMMARY 

Several basic grounding methodologies exist, and 
each has its merits. The resistance-grounded neutral sys- 
tem is superior for mining applications involving portable 
or mobile equipment. The design of ground beds is a 
complex field, and many variables must be examined in an 
attempt to derive an optimum configuration. A low value 
of resistance is of primary importance so dangerous poten- 
tials are not developed on machine frames. High potential 
gradients in the ground-bed area must also be avoided to 
prevent injury to personnel. A study of electric shock and 
its effects on humans is helpful in further delineating this 
subject. Formulas have been presented that may be used to 
predict the earth resistance of a particular metallic array 
or to determine how much buried metal is needed to 
achieve a desired value. In order to verify the ground-bed 
earth resistance, a description of ground test instrumen- 
tation, its utilization, and data interpretation was also 
included. When designing a ground bed, corrosion effects 
and soil-heating phenomena, caused by current flow in the 
ground system, must be considered. 

The resistivity of the soil in which the ground bed is 
immersed has a significant effect upon its earth resis- 
tance. Resistivity in turn is influenced by other factors 
such as earth composition, temperature, and moisture, 
and a thorough understanding of these relationships will 
be of use in metallic grounding-network design. Instru- 
mentation was again discussed, as well as practical appli- 
cations such as the determination of the best location for a 
ground bed. Chemical treatment of soils to increase con- 
ductivity and attenuate seasonal resistivity variations 
was reviewed. 

Correct selection and coordination of protective cir- 
cuitry is essential to gain the full benefits of a low- 
resistance ground bed. Protective circuitry must be in- 
stalled to monitor current flow in the ground conductors or 
the potential drop across the neutral grounding resistor. 
When properly coordinated, this protective circuitry will 
quickly shut down faulty sections of the electrical system. 

In the event of a fault or short circuit on a piece of 
mine machinery, its frame may become hot or elevated 
above ground potential. An unsuspecting miner could be 
seriously injured or killed if the machine is touched. 
Fast-acting relays and circuit breakers will minimize the 
length of time during which this shock hazard exists, and 
the bad circuit will be isolated from the remainder of the 
system. These protective devices form the subject of chap- 
ters 9 and 10. The grounding conductors that tie equip- 
ment frames to the safety ground bed are discussed in the 
next chapter, "Distribution." 



REFERENCES 

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5. Card, R. H. Earth Resistivity and Geological Structure. 
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181 



6. Clark, H. H., and C. M. Means. Suggested Safety Rules for 
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7. Clark, R. J., and B. 0. Watkins. Some Chemical Treatments 
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9. Cooley, W. L. Design Consideration Regarding Separation of 
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and Codification of Ground Bed Construction and Measurement 
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11. Dalziel, C. F. Effect of Wave Form on Let-Go Currents. 
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9, Feb. 1972. 

13. . The Threshold of Perception Currents. Trans. 

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14. Dalziel, C. F., J. B. Lagen, and J. L. Thurston. Electric 
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16. , Re-evaluation of Lethal Electric Currents. 

IEEE Trans. Ind. and Gen. Appl., v. 4, Sept./Oct. 1968. 

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Voltages. Trans. Am. Inst. Electr. Eng., Part 2, v. 75, May 1956. 

18. Dawalibi, F., and D. Mukhedkar. Optimum Design of 
Substation Grounding in a Two-Layer Earth Structure; Parts I, II, 
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1975. 

19. Dobrin, M. B. Introduction to Geophysical Prospecting. 
McGraw-Hill, 1952. 

20. Dornetto, L. D. The Importance of Grounding Systems in 
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June 1972. 

22. Dwight, H. B. Calculations of Resistances to Ground. Trans. 
Am. Inst. Electr. Eng., v. 55, Dec. 1936. 

23. Eaton, J. R. Grounding Electric Circuits Effectively; Parts 
I, H, and III. Gen. Electr. Rev., v. 44, June 1941. 

24. Fawssett, E., H. W. Grimmitt, G. F. Shorter, and H. G. 
Taylor. Practical Aspects of Earthing. J. Inst. Electr. Eng. (Lon- 
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25. Friedlander, G. D. Electricity in Hospitals: Elimination of 
Lethal Hazards. IEEE Spectrum, v. 8, Sept. 1971. 

26. Giao, T. N., and M. P. Sarma. Effect of a Two-Layer Earth 
on the Electric Field Near HVDC Ground Electrodes. IEEE Trans. 
Power Appar. and Syst., v. 91, Nov./Dec. 1972. 

27. Gienger, J. A. Fourteen Years of Data on the Operation of 
One Hundred Ungrounded 240 and 480 Volt Industrial Distribution 
Systems. IEEE Trans. Ind. and Gen. Appl., v. 2, Mar./Apr. 1966. 

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29. Gross, E. T. B., B. V. Chitnis, and L. J. Stratton. Grounding 
Grids for High-Voltage Stations. Trans. Am. Inst. Electr. Eng., 
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30. Hamilton, D. E. Mine Power Systems: What's Your Ground 
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32. Institute of Electrical and Electronics Engineers (New 
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33. Jacot, H. D. Grounding Practice in Coal Mines. Pres. at Min. 
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182 



CHAPTER 8.— DISTRIBUTION 



The distribution system within a mine consists of 
various types of cables that connect equipment to power 
supply, the conductors that form the trolley system used in 
many underground mines, and the overhead lines that 
distribute power in some surface mines. The character of 
the mining operation imposes constraints on the distribu- 
tion system unlike those of other industries and magnifies 
its importance within the overall power system. Mining is 
by definition constantly mobile; hence, the distribution 
system must be handled and extended frequently and can 
be susceptible to damage from mobile equipment. The 
mobility in turn necessitates efficient methods for joining 
cables and repairing them in order to minimize production 
downtime and operating costs. In all mines there is the 
potential for electric shock when handling distribution 
components. In the hazardous environment of an under- 
ground coal mine, damaged systems can be a potential fire 
and gas-ignition source. Proper installation and correct 
handling practices are essential if these hazards are to be 
minimized. 

This chapter's purpose is to introduce the various 
distribution components used in mine power systems, as 
well as to discuss their construction, installation, and 
maintenance. Cable systems are covered first and com- 
prise the majority of chapter content because of their 
uniqueness to mining. Typical trolley-system arrange- 
ments are then presented, and the chapter is concluded 
with a brief introduction to overhead lines. 



NATURE OF CABLE DISTRIBUTION 

It was shown in chapter 1 that cables can carry the 
electricity from the substation, where the power is taken 
from utility company lines, to the point of utilization by a 
mining machine, pump, conveyor belt, or other piece of 
equipment. There are many possible variations in mine 
distribution, and several types of cables can be put to a 
similar use. Only the most typical schemes are covered in 
this chapter, but some notable exceptions are included. 

Representative systems are depicted for underground 
coal mines in figure 8.1 and for surface coal mines in 
figure 8.2. Obviously, the circuits shown in the figures are 
only simplified examples of actual mine systems. In prac- 
tice, an underground coal mine would not have one long- 
wall, one continuous mining section, and one conventional 
section, but several continuous mining sections or several 
conventional sections in addition to one or more longwall 
units. Surface mines would usually have more than one 
dragline and one stripping shovel, not necessarily all 
electrically powered. 

As might be supposed, the kind of cable is tied to the 
application. Examination of figures 8.1 and 8.2 indicates 
that some cables remain in stationary locations for several 
years, while others are moved frequently. The cables that 
are connected to mining machines are termed portable by 
the Insulated Cable Engineers Association (ICEA) stan- 



1 The author wishes to thank Robert H. King, who prepared original 
material for many sections of this chapter. Thanks are also extended to 
James N. Tomlinson, who assembled the original section on splicing, and to 
George Luxbacher, who assembled the original material on conductor 
ampacities and cable derating. 



dards (19-2 1). 2 The Code of Federal Regulations uses the 
term trailing cables for the specific variety of portable 
cables used in a mine (38). Trailing cables are flame- 
resistant flexible cables or cords through which electrical 
energy is transmitted to a machine or accessory. 

In underground mines, trailing cables are generally 
attached to the inby end (toward the face) of the power 
center or distribution box. The portable cables that feed 
the power center or are attached to the outby end (toward 
the portal or shaft) have to be moved when the power 
center is advanced and retreated (perhaps once every 2 
weeks), but they are not moved as often as the trailing 
cables. The most stationary cables are those that bring 
power into the mine, for instance down the borehole and 
from the borehole to the portable switchhouses. These are 
the feeder cables. A special type, designated mine power 
feeder, can be used for installations that may not be moved 
for several years. However, the use of the word feeder here 
is to denote a cable type rather than a function in 
distribution. Both feeder and portable cables can be used 
for feeder applications, where the cable supplies two or 
more major loads (38). 



2 Italicized numbers in parentheses refer to items in the list of references 
at the end of this chapter. 



Entry, shaft, 
or borehole 



| I Switchhouse 



KEY 
Feeder or borehole cable 
Feeder cable 
Portable cable 
Trailing cable 




CONVENTIONAL UNIT 

4 Shuttle car 
4 Shuttle car 



4 Water pump 
4 Belt feeder 
4 Roof bolter 
4 Coal drill 
4 Loading machine 
4 Cutting machine 



CONTINUOUS UNIT 

4 Shuttle car 
4 Shuttle car 
4 Water pump 
4 Roof bolter 
4 Belt feeder 
4 Continuous miner 



LONGWALL UNIT 
4 Hydraulic pump 
4 Hydraulic pump 
4 Face conveyor 



Master 
control 



4 Face conveyor 
4 Stage loader 
4 Shearer 



Figure 8.1. —Cable distribution in underground coal mines. 



183 



KEY 
/ Feeder cable 

2 Portable cable 

3 Trailing cable 



Switchhouse 




3 Dragline 



1" 


2 


\ V 


3 

3 


Shovel 


, r 


2 


A V 


/ 


Switchhouse 




V- 


-o= 


Water pump 
Lighting 




1 l« 


3 



Power 
center 



Figure 8.2.— Cable distribution in surface coal mines. 



Similarly, in surface mines, the cables that feed from 
the switchhouses or unit substations to mobile equipment 
are trailing cables. Those moved only occasionally, which 
are not connected directly to a machine, are portable 
cables. Stationary (or near so) cables can be feeder or 
portable types. 

Moving the cable is a constant task both under and 
above ground. Some trailing cables are placed on reels or 
spools to facilitate moving. Prime instances of reeled 
cables are cables associated with the reeling devices on 
board shuttle cars and with mobile cable reels used in 
conjunction with many draglines. Trailing cables without 
reels are usually termed drag cables. Regardless of the 
application, cables are heavy and cumbersome and must 
often be manipulated by hand. Although the most fre- 
quent personnel injuries are strains, bruises, and frac- 
tures, cable handling is always potentially hazardous, and 
investigations in mines have indicated that exposed "live" 
conductors are a too-common occurrence. Indeed, most 
fatalities in cable-handling accidents are a result of rou- 
tine handling of unshielded cable (25). 

Constant handling also imposes considerable stress 
on the cables. While cable life is rated by manufacturers at 
up to 20 yr for other industrial applications, in an under- 
ground mine the actual cable life does not even approach 
this. Mine personnel have estimated the life of continuous 
miner cables at 8 months; roof bolter cables at 7 months; 
and shuttle car cables at 3 months (25), and within this 
lifespan the cable usually requires frequent repair. It has 
been estimated, for example, that 75% of the total ma- 
chine downtime for shuttle cars is cable related. 



CABLE COMPONENTS 

Cables are made up of three basic components: the 
conductor, the insulation, and the jacket, although there 



may also be fillers, binding, shielding and armor. In basic 
cable construction, the conductors are surrounded by in- 
sulation and the jacket covers the insulation. The design of 
these components is heavily dependent upon the physical 
stresses that the cable must withstand in the mine envi- 
ronment, including tension, heating, flexure, abrasion, 
and crushing. Hence, a discussion of typical stresses is 
helpful prior to describing component specifics, cable 
types (the various component assemblies into cables), and 
cable coding. 

High cable tensions are characteristic of both drag 
and reeled cables. When combined with other stresses 
such as flexure and twisting, tension can be very harmful 
to cable life. Drag cables are pulled around pillar corners, 
through mud, and over jagged rocks where the drag 
resistance is high. Consequently, a considerable force can 
be required to drag the cable and, thus, high tensions can 
develop. 

Machinery that utilizes cable-storage reels also fre- 
quently causes excessive cable tensions (13). For instance, 
the stored cable on the shuttle car is either payed out of 
the reel or spooled up into the reel as the machine is 
trammed. The tension required is dependent upon mine 
conditions, machine type, and cable size, but must be 
sufficiently high to prevent running over or pinching slack 
cable. However, if tensions become too high as a result of 
sudden jerks on the cable, cable and splice failures can 
become excessive. In addition, instantaneously high cable 
tensions can result in cable whipping. This is common 
with shuttle cars and also occurs on other machines that 
utilize cable reel storage devices, such as roof bolters, coal 
drills, and cutting machines. This whipping action is a 
hazard to mine personnel, who may be struck by the cables 
as they handle the cables or work nearby. 

In addition to excessive cable tensions, high cable 
temperatures frequently occur on machinery that utilizes 
cable-storage reels (6). The cable is wound on the reel, 
layer upon layer. Such layering prevents the cooling action 
of circulating airflow, and heating occurs. Consequently, 
the cable jacket and insulation may become softened and 
more susceptible to damage from cutting, tearing, and 
abrasion. If excessive temperatures occur, the cable jacket 
and insulation can actually blister or crack, becoming 
brittle. Thus, the physical damage caused by heating poses 
another hazard to mine workers who must handle the 
cable, especially in a wet mine environment. 

Another common cable stress prevalent in all mining 
cables is cable flexure. As with any material that is bent, 
internal tension and compression occur in flexed cables. 
These stresses cause relative movement of individual wire 
strands, abrading one wire against another and gradually 
deteriorating the conductors. Stresses fatigue the conduc- 
tors, making them brittle and more susceptible to further 
damage. 

Abrasion is also deleterious to cables and can have 
severe consequences. Cutting or tearing can occur when 
the cable becomes snagged or caught on rock, nails, and so 
on (6). Ripping or tearing of the cable jacket and insulation 
often results. Such damage can cause immediate cable 
failure, but more often than not, the damage goes unno- 
ticed. In a wet environment, water penetration can create 
a current path to the outer surface of the cable. An 
individual could come in contact with the wet cable 
several feet from a damaged area and still receive a shock 
that might be fatal. 

Another important cause of failure is cable crushing 
(6). This is usually the result of runovers or pinching the 



184 



cable with a machine frame. Here, the conductors are 
compressed against one another or against the machine, 
causing the insulation and jacket to split, as well as 
damaging the conductors. Even if there is no immediate 
failure, line-to-line or line-to-neutral faults that result in 
nuisance tripping of the protective circuit breakers can 
occur later. Water penetrating into damaged areas of the 
jacket can eventually work into areas of damaged insula- 
tion causing short circuits or a safety hazard. 

Conductors 

Line and ground currents are carried by either copper 
or aluminum conductors, depending on the specific char- 
acteristics required. Copper has high conductivity, is 
heavier and more flexible, but also more expensive. Be- 
cause of its greater flexibility, it is used in all portable 
mining cables. 

Copper cable conductors are usually composed of 
many fine wires combined into strands. Varying numbers 
of strands form the conductor. At the cable manufacturing 
plant, a cold-drawing process is used in which the copper 
rod passes through successively smaller dies to reduce its 
diameter (5). This process hardens the copper and makes it 
less flexible, so that if a soft-temper copper (strength about 
24,000 psi) is required, the wire must be annealed. Con- 
ductors that require a high tensile strength but are not 
bent frequently use medium- to hard-temper copper; 
medium-hard is rated at 40,000 psi. 

Copper conductors can become annealed in service if 
they are used at high operating temperatures for long 
periods of time. In fact, copper can lose 5% of its original 
tensile strength in 10,000 h at 70°C (5). Cable manufac- 
turers should always be consulted about the capability of 
their products to resist annealing when installed as bore- 
hole or high-tension overhead cables. To prevent corrosion 
by insulation vulcanizing agents, copper strands are usu- 
ally coated or tinned with lead or tin alloys, though this 
reduces the surface conductivity. 

Aluminum conductors are also used in mines. Alumi- 
num is cheaper, lighter, and less flexible, and has lower 
conductivity than copper. Aluminum conductivity is 61% 
that of copper; therefore, an aluminum conductor must 
have a cross-sectional area 1.59 times that of copper to 
have an equivalent dc resistance. However, copper conduc- 
tors weigh 3.3 times as much as aluminum; so even though 
the cross-sectional area of an aluminum conductor is 
greater, the total weight of an equivalent-resistance alu- 
minum conductor is less. Poor flexibility eliminates the 
use of aluminum in trailing cables. Aluminum is some- 
times used for feeder cables because of its lower cost, but 
problems can arise in jointing. An improperly constructed 
joint can allow the formation of aluminum oxides, which 
increase resistance and cause heating at the connection. 
Extreme care must also be taken to exclude moisture from 
any copper-to-aluminum joints because of the potential for 
electrolytic corrosion of the aluminum. 

Conductor Sizes 

The cross-sectional area of conductors is important for 
mechanical strength and is closely related to current- 
carrying capacity. Since the proper capacity is both a legal 
requirement and a desirable practice for safe operation, an 
understanding is needed of the methods commonly used to 
specify cross-sectional areas and ampacities. In the United 



States, both the American Wire Gauge (AWG) (or Brown 
and Sharpe Gauge) and circular-mil designations (MCM) 
are used (1 cmil is the area of a circle that is 1 mil in 
diameter). The AWG specifies 38 steps or sizes between No. 
36, which is 0.0050 in. in diameter, and No. 4/0, which is 
0.4600 in. in diameter (5). These sizes closely conform to 
the steps of the wire-drawing process. Table 8.1 specifies 
the cross-sectional areas and equivalent circular-mil sizes 
for some of the AWG designations. The 38 intermediate 
sizes are calculated in a geometric progression relating 
the ratio of any diameter to the next smaller or larger by: 



39 



fo. 



4600 



0050 



= 1.1229. 



(8.1) 



Table 8.1. —Conductor sizes and cross-sectional areas 



Conductor size 



Cross-sectional 
area 



cmil 



AWG: 

22 640 0.000502 

20 1,020 .000804 

19 1,290 .00101 

18 1,620 .00127 

17 2,050 .00161 

16 2,580 .00203 

15 3,260 .00255 

14 4,110 .00322 

13 5,180 .00408 

12 6,530 .00513 

11 8,230 .00646 

10 10,380 .00808 

9 13,090 .0102 

8 16,510 .0129 

7 20,820 .0164 

6 26,240 .0206 

5 33,090 .0260 

4 41,740 .0328 

3 52,620 .0413 

2 66,360 .0521 

1 83,690 .0657 

1/0 105,600 .0829 

2/0 133,100 .1045 

3/0 167,800 .1318 

4/0 211,600 .1662 



Conductor size 



Cross- 
sectional 
area, in 2 



MCM: 

250 0.197 

300 .236 

350 .274 

400 .315 

450 .353 

500 .392 

550 .432 

600 .471 

650 .510 

700 .550 

750 .589 

800 .628 

900 .707 

1,000 .786 

1,100 .863 

1,200 .942 

1,250 .981 

1,300 1.02 

1,400 1.10 

1,500 1.18 

1,600 1.26 

1,700 1.33 

1,750 1.37 

1,800 1.41 

1,900 1.49 

2,000 1.57 



Short-cut conductor-size approximations can be made 
by applying some simple rules if a table is not available. 
For example, the diameter will be doubled or halved by 
moving six sizes up or down the table. The weight, area, 
and dc resistance is doubled or halved by moving three 
gauge sizes, and they are changed by a factor of 10 over 10 
gauge sizes. A convenient reference point from which to 
apply these rules is the No. 10 wire since its diameter is 
about 0.1 in., its dc resistance is nearly 1 fl per 1,000 ft, 
and it weighs 10^ lb per 1,000 ft. In applying these rules, 
it should be remembered that the outer diameter and 
weight of conductors depends on the stranding configura- 
tion, which is described below. 

Federal regulations require grounding conductors to 
have at least one-half of the cross-sectional area of the 
power conductors (38). When the power conductor is a No. 
8 AWG or smaller, the grounding conductor should be the 
same size as the power conductor. The ground-check (pilot) 
conductor must not be smaller than a No. 10 AWG (38). 



185 



Conductor Stranding 

In order to obtain the required flexibility, mining 
cable conductors are made with numerous small wires 
rather than a single solid copper rod. These small wires 
are wound or laid together in strands, which are wound 
together in a rope in specific patterns. In a shuttle car 
cable, 37 wires are wound or bunched together, then 7 of 
these strands are spiraled together to form the conductor. 
Consequently, the total number of wires in this case is 259 
and equals the number of strands multiplied by the 
number of wires in each. The cross-sectional area of a 
stranded conductor is defined as the sum of the area of its 
component wires. 

In the simplest terms, conductor flexibility is greatest 
when the largest number of small-diameter wires is used. 
However, a certain amount of tensile strength is also 
required in mining cable conductors, and the tensile 
strength is greatest when a small number of larger wires 
is used. The design of a specific cable must therefore 
optimize these opposing factors, while taking into account 
the effects of twisting and bunching. Different applica- 
tions obviously necessitate different configurations. The 
engineer must examine cable stranding specifications 
carefully and select the one that best suits the application. 
Where historical information is not available, several 
types should be tried to find the best performer. Flexibility 
is also influenced by the method of insulating the power 
and ground-check conductors and applying the overall 
jacket. 

Insulation 

Insulation of mining cables is required to withstand 
stress from heat, voltage, and physical abuse. The insula- 
tion must be specially designed not only to protect mine 
personnel from electric shock, but also to separate power 
and grounding circuits effectively. 

Heating affects insulating materials in different 
ways, depending on their chemical composition. Heating 
either softens insulation, causing it to lose physical 
strength, or causes it to age or become brittle. Conse- 
quently, heat can make insulation lose its original shape, 
tensile strength, cut resistance, elongation, and effective- 
ness as an insulator. The main sources of heat are the 
environment, related to the ambient temperature, and 
power (PR) loss in the cable conductors. Hence, cable 
heating is directly connected to the maximum current the 
conductors can carry safely. 

Cable manufacturers usually prefer to use a thermo- 
setting insulation. After being extended over the conduc- 
tors, this insulation changes chemically by vulcanizing 
into a material that softens very little within the rated 
temperature range. The most common insulating com- 
pounds in this group are neoprene, styrene butadiene 
(SBR), ethylene propylene (EPR), and crosslinked polyeth- 
ylene (XLP). These compounds are usually mixed with 
other materials to achieve improved physical and electri- 
cal properties. 

SBR is used in 600-V trailing-cable insulation. It has 
a high modulus of elasticity, good flexibility, and a 75°C 
temperature rating, and resists damage by crushing from 
runovers and rock falls. EPR has replaced SBR in many 
trailing cables because it allows the cable rated voltage to 
be increased to 2,000 V and the temperature rating to 
90°C, while maintaining the same insulation thickness as 
SBR and neoprene. The EPR emergency-overload rating is 



130°C, and the short-circuit rating is 250°C. XLP is also 
rated at 90°C for normal operation and is used in high- 
voltage (> 1,000 V) mine-feeder and portable strip-mining 
cables. XLP is a rather stiff material, however, and is not 
recommended for reeling applications. 

The cable voltage rating is closely associated with the 
maximum anticipated operating voltage. The most com- 
mon ratings for mining cables are 600 V, 2 kV, 5 kV, 8 kV, 
15 kV, and 25 kV. The 5-kV, 8-kV, 15-kV, and 25-kV ratings 
are used primarily for stationary feeder cables and are 
generally not connected to mining equipment, except in 
surface mines. Usually, 4.16-kV distribution requires 5-kV 
rated cables, 7.2 kV requires 8 kV, and 12.47 kV and 13.2 
kV require 15-kV ratings. The utilization voltages of 250 
Vdc, 440 Vac, and 550 Vdc usually call for 600-V or 2-kV 
cables, and medium-voltage applications (661 to 1,000 V) 
need 2-kV insulation. 

The voltage rating of an insulation is actually based 
on its ability to withstand a test voltage that is many 
times the anticipated operating voltage, for a specified 
period of time. The test procedure and specifications are 
published in ICE A standards (19-21). Insulating com- 
pounds have different voltage ratings, which are usually 
expressed as the amount of voltage they can withstand per 
mil of thickness. Consequently, higher voltages can be 
used with any compound by increasing its thickness. 
Insulation thicknesses are also specified by ICEA. 

Insulation must resist damage from corona, particu- 
larly in high- voltage applications, as discussed in detail in 
chapter 17. The term partial discharge describes the type 
of corona stress imposed on cables. Partial discharges 
deteriorate insulation by ion bombardment and chemical 
action from ozone, nitrogen oxides, and nitric acid, which 
can occur in such voids as found between a stranded 
conductor surface and the insulation. Hence, insulation 
voids must be minimized and the insulation must resist 
the formation of this type of corona. ICEA standards 
specify corona-extinction voltage levels for insulation (19- 
21). 

Ozone resistance is important for high-voltage cable 
insulation and sometimes for low voltage, and standards 
are again given by ICEA. Ozone is formed when electrical 
discharge is present in air, and it attacks compounds 
containing double carbon bonds, by splitting the carbon 
chain and deteriorating the material. Radiating cracks are 
a physical symptom of this occurrence. 

Insulation must withstand cold temperatures as well 
as heat, particularly in surface operations: some of the 
open-pit iron mines in Minnesota and Michigan, for exam- 
ple, have experienced temperatures as low as -50°C. 
Cables stored on the surface at underground minesites are 
also exposed to extremely low temperatures. Most prob- 
lems occur when a cold cable must withstand mechanical 
stress, such as bending or impact. 

Cable Jacket 

The main purpose of the jacket is to provide protection 
for the inner components and hold the assembly in the 
designed configuration. Jackets are not required to pass 
ICEA voltage withstand or insulation resistance tests, but 
tests for tensile strength, elongation, and aging are man- 
datory. Ozone and discharge-resisting jackets must also 
pass surface-resistivity and partial-discharge tests. Min- 
ing cable jackets must withstand an extensive tempera- 
ture range, maintaining their physical properties through- 
out, and furthermore, they must not deteriorate when 



186 



exposed to direct sunlight. Obviously, resistance to abra- 
sion, crushing, tearing, and impact are extremely impor- 
tant. Cable jackets must also be resistant to the chemical 
action of acid or basic mine water and hydraulic fluids, 
and underground coal mine cable jackets must be flame 
resistant. Finally, jackets must exclude moisture and be 
very flexible. 

One of the most commonly used materials for cable 
jackets is neoprene, a chloroprene polymer. Nitrile buta- 
diene and polyvinyl chloride (NBR/PVC) is also used, 
particularly where jacket coloring is desired. Chlorosul- 
fonated polyethylene (CSP) or Hypalon synthetic rubber is 
also used extensively, especially in combination with 90°C 
EPR insulation. EPR is used where extreme cold is en- 
countered and flame resistance is not essential. Armored 
cables are used in some borehole applications. Here the 
jacket is a heavy metallic covering that affords extra 
protection to the conductors. 

Cable Shielding 

The ICEA defines the practice of shielding an electri- 
cal power cable as confining the electric field to the inside 
of the cable insulation or assembly with a grounded 
conducting medium called a shield (19-21). Two shield 
types are used in practice: the conductor shield and the 
insulation shield. Shown in figure 8.3, the conductor 
shield is placed between the conductor and the insulation, 
and the insulation shield surrounds the insulation. 

Two distinct types of materials are employed in con- 
structing cable shields: nonmetallic and metallic. Nonme- 
tallic shields may consist of a conducting tape or a layer of 
extruded conducting compound. The tape may be made 
from conducting compound, be a conducting fibrous tape, 
or be a fibrous tape faced or filled with conducting com- 
pound. A typical conducting compound is carbon- 
impregnated rubber, which is commonly referred to as a 
conductive-rubber, semiconducting, or semicon shield. Me- 
tallic shields are nonmagnetic and may consist of a thin 
metal tape, wire-woven braid, or concentric serving of 
wires. Copper-braided shields may be made entirely of 
copper wires or have nylon twine in combination with 



Insulation 



Conductor 
shield 




Conductor 



Conductor 



copper wires. Nonmetallic and metallic elements may be 
juxtaposed to form the shield. 

Conductor shields are made of nonmetallic materials 
and are used only in high-voltage cable. The roles of this 
shield type are to eliminate air spaces or voids between the 
conductor and the insulation and to present a smooth 
electrode to the inner insulation surface. To be effective, it 
must adhere to or remain in intimate contact with the 
insulation under all conditions. This can substantially 
reduce the number of sites where partial discharge can 
form and helps reduce electrical stress on the insulation by 
uniformly distributing the electrical field about the con- 
ductor. The use of conductor shields becomes critical at 
higher operating voltages, especially 12.47 kV and above. 

Insulation shields can perform three principal func- 
tions. If placed directly over individual conductor insula- 
tions, along with confining the electric field caused by 
conductor current within the insulation, the shield helps 
to maintain a symmetrical radial distribution of voltage 
stress within the dielectric. The possibility of partial 
discharges is minimized by precluding tangential and 
longitudinal stresses, and insulation is utilized to its 
greatest efficiency and in the direction of highest strength. 
This again becomes critical at higher operating voltages. 
Insulator shields also provide a continuous capacitance to 
ground for the conductor along its entire length. The 
uniformity is important in terms of transients on the 
power system, and this is discussed in chapter 11. 

The third function of insulation shields is the most 
important for mining in view of the extensive handling of 
cables: reducing the hazard of electric shock. A major 
cause of electrical fatalities in mining has been workers' 
cutting into energized unshielded cables, for instance, 
during repair. Another source has been handling of ener- 
gized unshielded cables with damaged jacketing and insu- 
lation or splices (the spot where a cable has been repaired). 
An insulation shield can be thought of as a safety barrier 
to penetrating metallic objects. If the percent of coverage 
of the shield over the insulation is high enough and its 
impedance is low enough, any metallic object compromis- 
ing the conductor insulation will establish a fault between 
the power conductor and the grounded shield, with suffi- 
cient current to trip the ground-fault protective circuitry. 
Damage to insulation and jacketing, such as a pinhole, 
that would cause a handling danger to unshielded cable 
also creates a probable ground fault in cables with insu- 
lation shields. An individual touching the penetrating 
metallic object or handling the damaged shielded cable 
should be safe from electrocution. 

Insulation shields are usually metallic. Recently, how- 
ever, semicon insulation shields for trailing cables have 
found application in the United Kingdom, Australia, and 
to a lesser extent, the United States. This is to take 
advantage of semicon flexibility, especially in reeled-cable 
situations. 



Insulation 


#5|r 


Conductor 




«titi_ 


_ Insulation 
shield 




Insulation 




Figure 8.3.— Shield types. 



CABLE TYPES 

An identifying code, related to standard specifications 
designated by ICEA, is embossed on the cable throughout 
its entire length. The code includes any approval number 
for flame resistance by the Mine Safety and Health 
Administration (MSHA) and approval by the Common- 
wealth of Pennsylvania (indicated by the letter P preced- 
ing the MSHA approval number). MSHA approval is 
mandated for cables in underground coal mines, and the 



187 



Pennsylvania approval is necessary for cables used in 
underground coal mines in that State. 

The code includes the term nJc where n is the number of 
power conductors in the cable, an approved voltage designa- 
tion, and letters describing the cable type. Table 8.2 summa- 
rizes the meaning of the letters used in the code, and table 
8.3 presents the codes for typical cable types used in mining. 
Figures 8.4, 8.5, and 8.6 correspond to table 8.3 for un- 
shielded round, unshielded flat, and shielded cable configu- 
rations, respectively, and detail the cable components as seen 
in cross section. Photographs of actual mining cables are 
provided in figures 8.7, 8.8, and 8.9 and show both side and 
cross-sectional views. Figures 8.10 and 8.11 are similar to 
figures 8.1. and 8.2 and show common applications of cable 
types in mine power systems. 

Figure 8. 7 A is a single-conductor cable insulated for 
use at 600 V. This specific cable is not widely used. 
However, it has found application on twin-reel dc shuttle 
cars and small locomotives with reels; therefore, it must be 



highly flexible. Single-conductor cable similar to that 
shown is used extensively for connections inside power 
equipment, and a typical voltage rating is 15 kV for 
system voltages less than that level. 

The most common dc shuttle-car cables are types W 
and G, figures 8.8A and 8.8.B, respectively. The flat con- 
figuration is used since it allows an increased length on 
cable reels and is less susceptible to runover damage than 
round cables. The type W is used where diode grounding is 
allowed in lieu of a separate grounding conductor. Because 
shuttle car cables are damaged frequently, type W is 
preferred by some mine operators since it is easier to 
repair (splice). 

Flat cable types employed for ac shuttle cars are 
shown in figure 8.8C and 8.8D. The three power conduc- 
tors are separated by two grounding conductors in figure 
8.8C and by one grounding and one ground-check conduc- 
tor in figure 8.8D. These cables are also used on other 
equipment with reels, such as cutting machines and drills. 



Table 8.2.— Letters used in alphabetic cable code 



Code Meaning Comments 

G Contains uninsulated grounding conductor(s) Common on low-voltage ac systems but used on dc systems where 

grounding conductors are needed. 

W Without uninsulated grounding conductor(s) Typical on dc diode-grounded systems but 1 insulated power conductor 

may be used as a grounding conductor. 

GC Includes insulated ground-check (pilot) conductor. Used where pilot-type ground-continuity monitoring is required, usually 

replaces 1 grounding conductor of type G cable. 

SH Shielded cable None. 

D Multiple insulation shields Shields surround each individual power-conductor insulation. 

C 1 insulation shield 1 shield surrounds entire cable assembly just inside jacketing. 

MP Mine power feeder None. 



Table 8.3.— Codes for typical cables used in mining. 



Code Components 

W Contains 2, 3, or 4 insulated power conductors 

G Contains 2 or 3 insulated power conductors and 1 to 3 

uninsulated grounding conductors. 
G-GC Contains 3 insulated power conductors, 1 or 2 uninsulated 

grounding conductors, and 1 insulated ground-check conductor. 

G + GC Contains 3 insulated power conductors, 3 uninsulated grounding 

conductors, and 1 insulated ground-check conductor. 

SH-D Contains 3 shielded insulated power conductors, 2 or 3 

uninsulated grounding conductors. 

SH-C Contains 3 insulated power conductors, 2 or 3 uninsulated 

grounding conductors, assembly shielded. 

SHD-GC Contains 3 shielded insulated power conductors, 1 or 2 

uninsulated grounding conductors, and 1 insulated 

ground-check conductor. 
SHD + GC Contains 3 shielded insulated power conductors, 3 uninsulated 

grounding conductors, and 1 insulated ground-check conductor. 
SHC-GC Contains 3 insulated power conductors, 1 or 2 uninsulated 

grounding conductors, 1 ground-check conductor, assembly 

shielded. 
MPF Contains 3 shielded insulated power conductors, 3 uninsulated 

grounding conductors. 
MP-GC Contains 3 shielded insulated power conductors, 2 uninsulated 

grounding conductors, and 1 ground-check conductor. 
1 Although not presently available, 2/C cable design for dc systems is possible. 



Comments 



See table 8.2. Flat or round cross section. 

Grounding conductors are placed in the interstices between the 

power conductors. Flat or round cross section. 
Ground-check conductor replaces 1 grounding conductor of type 

G cable. Flat or round cross section. Presently, for ac systems 

only. 1 
Similar to round 3/C type G cable but has ground-check 

conductor in cable center. 
Insulation shields about each individual conductor, grounding 

conductors contact shields. High-voltage cables usually have 

conductor shields. Round or flat cross sections. Presently for ac 

systems only. 1 A flexible portable cable. 
Shielding encloses all conductors and is located just under . 

jacketing. Grounding conductors should contact shield. Round 

or flat cross sections. Presently for ac systems only. 1 A flexible 

portable cable. 
Ground-check conductor replaces 1 grounding conductor of type 

SH-D cables. Round or flat cross section. Presently for ac 

systems only. 1 A flexible portable cable. 
Similar to round 3/C type SH-D cable but has ground-check 

conductor in cable center. 
Ground-check conductor replaces 1 grounding conductor of type 

SH-C cables. Round or flat cross section. Presently for ac 

systems only. 1 A flexible portable cable. 
Similar to round SH-D cable. Designed for relatively stationary 

high-voltage feeder applications. 
Similar to round SHD-GC cable. Designed for relatively 

stationary high-voltage feeder applications. 



188 




Insulated 

ground -check 

conductor 




Type W 



Type G 



Type G-GC 



Uninsulated 
grounding 
conductors 



Insulated 

ground -check 

conductor 




Type G+GC 



Figure 8.4.— Cross sections of round unshielded mining cables. 




Fillers, may not be needed if 
conductor insulation fills voids 



Jacket 



2/C type W 




2/C type G 




3/C type G 



Figure 8.5.— Cross sections of flat unshielded mining cables. 



3/C type G-GC 



Conductor 
insulation 



Conductor shield, 
copper braid if 
SHD-GC, metallic 
tape if MP-GC 




Grounding conductor, 
contacts shield 



Jacket 



Type SHD-GC or MP-GC 



Grounding conductor 
may contact shield 




Type SH-C 



Conductor shield 



Grounding conductors 
contact shield 




Jacket 



Flat type SH-D 



Jacket 



, Grounding conductors 
may contact shield 




Flat type SHC-GC 



Figure 8.6.— Cross sections of some shielded mining cables. 



189 




(A) 2/C typeW, 600V 



(A) 1/C.600V 




(B) 2/C type G, 600 V 





(B) 3/C type G-GC, 2,000 V 




(C) 3/C type G, 600 V 






(C) 3/C type G+GC, 2,000 V 

Figure 8.7.— Round unshielded mining cable. (Courtesy 
Anaconda Ericsson Co.) 



(D) 3/C type G-GC, 600 V 

Figure 8.8.— Flat unshielded mining cables. (Courtesy Anacon- 
da Ericsson Co.) 





(A) 3/C type SHD-GC, 2,000 V 




(C) 3/C type SHD-GC, 15 kV 







(B) 3/C type SHD+GC, 2,000 V (D) 3/C typeMP-GC, 15 kV 

Figure 8.9.— Round shielded mining cables. (Courtesy Anaconda Ericsson Co.) 



190 



KEY 
Borehole cable: 3/Ctype MP-GC; 5, 8, 15, or 25 kV 
3/C type MP-GC, SHD-GC, or SHD+GC; 5,8, 15, or 25kV 
3/C type SHD-GC or SHDtGC; 5,8, !5,or25kV 
3/C type G,G-GC,orG + GC; 2kV 
3/C typeGorG-GC, flat, 2kV 
2/Ctype WorG, flat, 2 kV 



Entry, shaft, 
or borehole 



I Switchhouse 

k 

Switchhouse 




CONVENTIONAL UNIT 

250-Vdc Shuttle car 
250-Vdc Shuttle car 

550-Vac Water pump 
550 -Vac Belt feeder 
550-Vac Roof bolter 
550-Vac Coal drill 
550-Vac Loading machine 
550-Vac Cutting machine 



CONTINUOUS UNIT (550 Vac) 

Shuttle car 
Shuttle car 
Water pump 
Roof bolter 
Belt feeder 
Continuous miner 



LONGWALL UNIT (550 Vac) 
4 Hydraulic pump 
4 Hydraulic pump 
4 Face conveyor 



^<L 



I Master 
I control 



4 Face conveyor 
4 Stage loader 
4 Shearer 



Figure 8.10.— Cable types for typical distribution systems in 
underground coal mines. 



The ac shuttle cars also utilize round cables of type G 
or type G-GC. The grounding conductors are placed in the 
interstices between the power conductors in the type G, 
and a ground-check conductor replaces one of the ground- 
ing conductors in the type G-GC (fig. 8.7B). In addition to 
limited use on shuttle cars, the majority of longwall 
shearer, face-conveyor, stage-loader, roof-bolter, feeder, and 
continuous-miner cables are of this type. In some in- 
stances, the G-GC configuration can initiate induced 
voltages in the frame-grounding system (see chapter 17). 
Therefore, the G + GC type shown in figure 8.7C was 
constructed. Here the three grounding conductors are laid 
symmetrically in each interstice, and the ground-check 
conductor is placed in the center of the cable. 

There are two basic configurations for shielded cables: 
the SH-D and the SH-C. As shown in figure 8.6, the 
shield of the SH-D cable surrounds each insulated conduc- 
tor; in the SH-C cable, the shielding encloses all power 
conductors and grounding conductors. The SH-D shield- 
ing is preferred because the grounding conductor is in 
intimate contact with the shield, and line-to-line leakage 
current is detectable since the shield surrounds each 
individual power conductor. The SH-C shield, a single 
braid over the entire assembly, is sometimes found in 
low- voltage and medium-voltage portable cables. However, 
special designs are required to assure consistent, low- 



KEY 

/ 3/C type MP, MP-GC, or MP + GC; 5,8, 15, or 25 kV 

2 3/C type SHD, SHD-GC, or SHD+GC, 5,8,l5,or 25kV 

3 3/C type SHD, SHD-GC, or SHD + GC, 2 kV 



2 Dragline 




2 Shovel 



3 Water pump 
3 Lighting 



Power center 



Figure 8.11.— Cable types for typical distribution systems in 
surface coal mines. 



resistance contact between the shield and the grounding 
conductors. 

In high-voltage cables, the insulation shield is gener- 
ally comprised of two parts: an extruded layer or wrap of 
semiconducting material applied directly over the insula- 
tion, and a metallic cover applied over the semiconducting 
layer. The semiconductive material is considered to have a 
100% coverage, but an associated high resistivity. If the 
metallic layer is composed entirely of copper braid, its 
coverage is 84% while the combination copper-nylon braid 
covers 60%. Shielding of unidirectional spirally wound 
wires, which gives 60% coverage, may also be used. The 
high- voltage insulation shield must be in intimate contact 
with the insulation under all conditions in order to be 
effective, and the metallic portion serves as a current- 
carrying medium for charging and leakage currents. Fed- 
eral regulations require SH-D shielding for high-voltage 
cables in underground coal mines. Both SH-D and SH-C 
shielding are permitted for medium-voltage cables. 
Medium-voltage cables used on reels do not have to be 
shielded if the insulation is rated at 2 kV (38). 

Two round shielded-cable configurations, SHD-GC 
and SHD + GC, are also used extensively for medium- 
voltage and high-voltage cables. The 2,000-V-rated SHD- 
GC cable, shown in figure 8.9A, and the SHD + GC cable 
in figure 8.9B are common on such equipment as 950-Vac 
continuous miners and longwall shearers, and on low- 
voltage surface coal mine equipment. Some high-voltage 
cables are required to be flexible, for example, surface 
mine shovel and dragline cables and underground mine 
distribution cables, which are connected to a portable 
power center. The SHD-GC cable shown in figure 8.9C is 
intended for this application. It is rated at 2, 5, 8, 15, or 25 
kV depending on insulation thickness. 



191 



Stationary power cables are often mine power feeders 
of the MP-GC type as shown in figure 8.9D (see also tables 
8.2 and 8.3). These cables can also be rated at 5, 8, 15, or 
25 kV, but they are less flexible and have higher tensile 
strength than the SHD-GC type. Shielding is similar but 
uses different materials. MP-GC cables are also designed 
to be used in boreholes, aerial installations, ducts, and 
direct burial. 

These are the basic power-cable types used currently 
in the mining industry. Other configurations are made for 
specific applications. For example, one double-drum shear- 
ing machine model requires a six-conductor cable with two 
ground-check conductors and a grounding conductor. Ca- 
ble manufacturers are usually willing to produce these 
special cables, but they are not a part of normal product 
lines and the possible variations are too numerous to 
include here. 



CABLE TERMINATIONS 

The termination or end of any cable must encompass 
a means of sealing and protecting the cable from the 
weather above ground and contaminants such as dust 
below ground. It must often provide a means of electrical 
connection with other conductors. Particularly in the case 
of high-voltage cables, considerable stress occurs on the 
dielectric between the terminating point of the cable 
shield, which is at ground potential, and the end of the 
conductor, which is at line potential. These electrical 
stresses are ameliorated through use of a stress cone that 
forms part of the termination device. 

The terminating device may take many forms, may be 
of varied complexity, and may be constructed from differ- 
ent insulating materials, depending on the cable type and 
the application. Taped terminations are very common, 
particularly at 15 kV and below. A simple sealing lug 
applied with insulating tape can be used on nonshielded 
cables, but where the cable is shielded, a stress-relief cone 
must be included. This may be preformed of rubber-like 
synthetic polymer and include an upper insulated cap, or 
may merely consist of lapped tape built up to the required 
cone shape. In either case, additional cover tapes are 
applied over the assembly and a rain hood or other 
protective housing may be added. An armor terminator 
provides a watertight grounding for armored cables and 
may be used in addition to a stress cone and insulation. 

A pothead is a form of termination housing used 
frequently in surface mines and above ground at under- 
ground mines. The pothead is hermatically sealed and 
thus provides maximum cable protection from the envi- 
ronment. A typical pothead for a shielded cable is shown 
in figure 8.12. Note that with shielded cables the termi- 
nation is taped prior to insertion in the pothead. In 15-kV 
applications and above, heated liquefied asphaltic or res- 
inous material is then poured into the pothead cavity. The 
rate of cooling of this dielectric material must be con- 
trolled to prevent the formation of voids. The pothead may 
include a number of aerial and cable connectors. 

Even though potheads are used, the standard 
termination-connector in mines is the coupler. An entire 
range of complex couplers has been developed specifically for 
the mining industry to accommodate the unique combina- 
tion of environmental factors and operating procedures. 



CABLE COUPLERS 

Couplers are the complex sophisticated plugs and 
sockets used throughout the mine distribution system to 
connect mobile machinery to trailing cables, to connect 
cables with one another, and to connect cables to power 
centers, switchhouses, and substations. All couplers have 
certain common characteristics: 

• They have either male contacts (plugs) or female 
contacts (sockets), 

• They are either line mounted (at the end of a cable) 
or gear mounted (located on a piece of equipment), 

• They are available in a wide voltage range, from 
high voltage (feeder cables) to low voltage (equipment 
related), 

• They are available in a range of sizes to accommo- 
date different types and ampacities of cable, 

• They all have grounding contacts and may also 
include ground-check contacts, 

• They all have sealing and locking devices and dust 
covers to protect the contacts when they are not in use. 

The complexity of couplers is a direct result of the 
mine environment in which they are used; they must 
resist damage, be sturdy enough to withstand repeated 
use, prevent electrical hazards, be watertight, be dust 
proof, and withstand heat and cold. Some models are rated 
explosion proof. The plugging mechanism must be easy to 
use yet secure. 

High-voltage couplers have been used in mine distri- 
bution for about 40 yr. Most of the initial problems 
encountered in 4,160- and 7,200-V systems have been 
resolved over the years through constantly improved de- 
sign. Operating failures are no longer common at these 
levels. However, some problems are still found in the 
15-kV class of couplers, and these have inhibited the 
switch to higher voltages by many mine operators. No 
ideal material has yet been developed for insulation; those 



Gasket 



Ceramic insulator 



Metal lid 



Grounding stud 



Stress- relief 
cone 

Grounding 
conductor 



Metal housing 




Terminal 
True seal fitting 
Hood nut 

Ferrule 



Cavity may be filled 
with dielectric material 

Cable support 



Figure 8.12.— Cable terminations for applications up to 15 
kV (all or a portion of cable weight is supported by pothead). 



192 



with excellent electrical and chemical properties have 
been found to have mechanical inadequacies, and vice 
versa. The combination of dust and dirt with high humid- 
ity and moisture found in underground mines has posed 
many problems. In too many instances, these difficulties 
have been compounded by neglect, impatience, and total 
disregard for the purpose of a component by those who use 
them (7). 

Coupler Contacts 

The general requirements for coupler contacts are 
summarized as follows. 

The coupler contact system should have 

1. Adequate current-carrying capacity and low resis- 
tance, 

2. The ability to withstand repeated coupling, 

3. Protection from worker abuse, 

4. A reliable and easy-to-make connection to the cable 
conductor, 

5. Oxidation and corrosion resistance, 

6. Uncoupling feature that allows a pilot or ground 
check to disengage first and the ground wires to uncouple 
last, 

7. A guidance system to prevent misalignment and 
bending during coupling, 

8. A feature to allow replacement of bent or damaged 
contacts. 

It is important that the male and female pins that mate as 
the coupler is connected are of adequate size and have low 
contact resistance to prevent excessive heating when car- 
rying current. 

Frequent coupling and uncoupling can lead to a poor 
contact, particularly when a coupler is dropped, not an 
infrequent occurrence. Contacts can be bent and become 
dirty. Poor alignment during coupling and attempting to 
force a connection can also bend the contacts. In either 
case, the resulting high-resistance connection can lead to 
problems with overheating. Coupler manufacturers have 
attempted to reduce damage to contacts by recessing them 
in the housing and adding guidance systems to facilitate 
alignment when coupling must be carried out in restricted 
spaces. 

Another possible failure point is the connection be- 
tween the cable conductor and the contact. Set screws, 
soldering, thermit welding, and brazing are various meth- 
ods for securing this connection. Extreme care must be 
taken when brazing or soldering these connections to 
remove excess flux, which can destroy coupler insulation. 
Severe vibration caused by dropping or by bouncing and 
bumping on a mobile machine such as a battery scoop can 
loosen a screw or crack a weld. The high-resistance broken 
connection then heats, which can cause insulation deteri- 
oration and a fault. 

Electrical voids and protrusions caused by an im- 
proper mating have great significance at voltages greater 
than 8 kV because these localized nonconformities can 
become partial-discharge inception points. Hence, the in- 
sulation should be made of corona-resistant materials and 
the contact design should minimize the occurrence of voids 
and protrusions. Some low-voltage couplers, for example, 
have a "self- wiping" action to improve the contact; other, 
high-voltage contacts employ a Multilam band for the 
same purpose. 






Coupler Insulation 

The general requirements for coupler insulation are 
as follows. 

The coupler insulation system should have 

1. Adequate dielectric strength, 

2. Adequate corona-extinction level, 

3. Adequate tracking resistance, 

4. Stress-relief feature, 

5. Adequate impulse level, 

6. Flame resistance, 

7. Resistance to moisture penetration, 

8. Insulators that align easily for coupling, 

9. Resistance to cracking, chipping, and bending, 

10. Resistance to heat deterioration, 

11. The ability to withstand repeated coupling, 

12. A feature that discourages phase reversal during 
mounting and coupling. 

To ensure that coupler insulation does not break down in 
normal service, it should have a dielectric strength equal 
to or greater than that of the cable entering the connec- 
tion. For high-voltage installations, the surface of the 
insulation should resist arc tracking, a process in which 
high-current arc discharges cross the insulator surface and 
carbonize the material, forming a conductive track. Keep- 
ing high-voltage insulators clean and dry will reduce the 
incidence of arc tracking. A common cause of moisture 
contamination is dropping the coupler on a wet mine floor. 
Insulation, particularly if it has been weakened by 
partial discharges, is subject to breakdown by high- 
impulse voltages called transients, which usually occur 
during switching. Insulation materials must be able to 
withstand repeated occurrences of these high voltages. 

Coupler Housing 

Characteristics required for the housing are as follows. 
The outer covering should have 

1. A reliable easy-to-make ground wire connection, 

2. A cable strain-relief mechanism, 

3. A guidance system that improves the ease of 
alignment for coupling, 

4. A durable material composition, 

5. The ability to withstand repeated coupling, 

6. Corrosion resistance, 

7. Grommet or packing gland of the correct size for the 
cable used, 

8. As little weight as possible, 

9. A feature that facilitates ease in coupling and 
uncoupling. 

If the coupler is classified as explosion proof, it incorpo- 
rates a packing gland at the entrance to the housing that 
usually consists of asbestos fiber packed tightly between 
the cable and bushing. To be rated explosion proof by 
MSHA, an explosion that occurs inside the shell should 
not ignite any methane-air mixture surrounding the cou- 
pler. Explosion-proof couplers are allowed inby the last 
open crosscut in underground coal mines by all State and 
Federal regulations. Connectors without packing glands 
can be used inby the last open crosscut if they have a pilot 
or ground-check circuit that interrupts the power before 






193 



the housing is opened. Instead of packing, non-explosion- 
proof couplers have a rubber grommet that allows cables of 
different diameters to fit into the same housing. 

The cable strain-relief clamp is located on the outside 
of the cable entrance and prevents cable tension from 
pulling the conductors out of their connections. The clamp 
may be drawn down on the cable jacket by tightening a 
bolt on either side. If the bolts are not sufficiently tight, 
the clamp will not prevent tensile pullout, and if too tight, 
the clamp will damage the cable insulation. 

Both packing glands and strain-relief clamps are 
made to fit a single cable jacket size or a small range of 
sizes. Thus knowledge of cable outer dimensions is neces- 
sary to match the coupler cable entrance to the cable. 
Tables 8.4 and 8.5 contain typical dimensions for round 
and flat cables, respectively (38). Variations in these 
values are allowed as long as the packing gland or strain 
relief is used. 

High-Voltage Couplers 

Couplers in the 15-kV, 500-A range are used as 
connections to switchhouses and power centers, to join 
high-voltage cables, and for high-voltage machines. A 
typical high-voltage coupler is shown in figure 8.13. In the 
following paragraphs, the numbers in parentheses refer to 
this diagram. 

A high-voltage coupler accommodates the three power 
conductors (4), one or more grounding conductors (14), and 
one or more ground-check conductors (15). The contacts (8, 
10, 11) are soldered and taped to the prepared conductor 
cables during installation. The contacts may be of copper, 
copper berylium, or in some cases, aluminum or brass. Male 
contacts have a split-pin design or incorporate a Multilam 
band of torsion-sprung louvres to improve the power contact. 



Table 8.4.— Typical diameters for round portable power cables 
in inches, 601 to 5,000 V 



G-GC, 

Conductor size „ . v 

AWG: 

4 1.25 

3 1.40 

2 1.48 

1 1.55 

1/0 1.74 

2/0 1.84 

3/0 1.99 

4/0 2.12 

MCM: 

250 2.30 

350 2.75 

1 Cable not made. 



SHC-GC, 
2kV 



SHD-GC, 
<3kV 



SHD-GC, 
3-5 kV 



1.39 


1.62 


1.78 


1.55 


1.77 


1.90 


1.62 


1.84 


1.98 


1.71 


1.92 


2.09 


1.89 


2.04 


2.18 


2.02 


2.18 


2.34 


2.16 


2.29 


2.46 


2.30 


2.45 


2.62 


2.48 


2.62 


2.76 


2.97 


( 1 ) 


( 1 ) 



Table 8.5.— Typical dimensions for flat portable cables in 
inches, 600 V 





Conductor 




2-conductor 




3-conductor, G 




W 




G 






AWG 


Major 


Minor 


Major 


Minor 


Major Minor 






axis 


axis 


axis 


axis 


axis axis 


8.... 




0.84 


0.51 


— 


— 


— — 


6.... 




.93 


.56 


1.02 


0.56 


1.65 0.67 


4.... 




1.05 


.61 


1.15 


.61 


1.85 .75 


3.... 




1.14 


.68 


1.26 


.68 


1.99 .77 


2.... 




1.24 


.73 


1.35 


.73 


2.10 .81 


1.... 




1.40 


.81 


1.55 


.81 


2.43 .97 


1/0. 




1.51 


.93 


1.67 


.93 


— — 


2/0. 




1.63 


.99 


1.85 


.99 


— — 


3/0. 




1.77 


1.03 


2.00 


1.03 


— — 


4/0. 




1.89 


1.10 


2.10 


1.10 


— — 


NOTE.— Dash indicates cable is not made. 








KEY 

/ Cable strain-relief clamp 8 

2 Packing-gland bushing g 

3 Plugged holes for pouring potting compound 10 

4 Power conductor (insulation wrapped) with // 
shielding tape for high voltage ]2 

5 Molded stress-relief cone (for high voltage) /J 

6 Shell engagement mechanism 14 

7 Insulation mounting flange /5 



Power-conductor contact 
Power- conductor insulation tube 
Grounding-conductor contact 
Pilot-conductor contact 
Metal lie -shell grounding point 
Coupler shell 

Grounding conductor (from cable) 
Ground -check conductor 



Figure 8.13.— Coupler components. 



194 



The insulation materials and configuration vary ac- 
cording to the manufacturer, but they commonly have 
three main parts: a molded stress-relief cone (5), insulation 
tubes (9), and a flange (7). The molded stress-relief cone is 
now tending to replace hand taping as a method of 
providing termination stress relief. It combines the func- 
tions of a stress-relief cone and seal and also serves to 
position the conductors. The insulating tubes push onto 
the tapered cylinders of the molding and encase the 
contacts. Resistant rubber-like polymer tubes are now 
finding favor over flexible rubber tubes or cups that have 
a tendency to fold when a misaligned coupling is at- 
tempted. Both types replace an earlier polyester insulator 
that could crack and chip under the abuse almost inevita- 
ble when coupling in the confined spaces of an under- 
ground mine. The insulation tubes attach to a rigid 
insulation flange that positions the assembly correctly, 
seals the contact area from the rest of the coupler, and 
attaches it to the housing. 

When the coupler assembly is complete, a potting 
compound may be poured into the coupler (3) to guard 
against the formation of moisture. The compound is de- 
rived from tung oil and sets to a gel-like consistency. 
Potting compound is not required for 15-kV couplers that 
use filler moldings, but is frequently used as an added 
precaution. Asphaltic compounds were originally used as 
coupler fillers but these were very difficult to remove if 
components were to be reused. 

The coupler housing (13) is metal, usually a high- 
strength, light-weight, corrosion-resistant cast aluminum 
that resists physical abuse yet is portable. The coupler 
housing incorporates a threaded collar or lock ring (6) that 
secures one coupler to its mate. Some designs have a 
pin-and-slot mechanism to reduce the number of turns 
required to lock the connection and simplify alignment. 



Low-Voltage Couplers 

The standard sizes for low-voltage and medium- 
voltage couplers are 225, 400, 600, 800, and 1,200 A. Their 
primary use is to connect mobile equipment to power 
centers and junction boxes, and to connect cables in the 
600- to 1,000-V range. Their construction is sturdy but less 
complex than that of high-voltage couplers. They have 
either a boxlike shape and are locked by a latch mecha- 
nism or a cylindrical lock ring similar to those on high- 
voltage couplers. They do not have stress-relief cones or 
packing compound. The packing gland is usually replaced 
by a rubber grommet seal, but these couplers do include a 
cable strain-relief clamp. Many different contact configu- 
rations are available to accommodate a wide range of 
equipment types. Lower powered couplers specialize in 
quick and easy connection and disconnection for equip- 
ment that must be changed out frequently. 



CABLE SELECTION 

The cable manufacturer can provide a proper cable to 
a mining company only if the exact operating conditions 
for the cable are specified. The purchaser has the respon- 
sibility for writing a purchasing specification that com- 
pletely describes the operating environment. A revised 
ICEA listing of the information to be supplied by the 



purchaser, given below, will be used here to describe the 
step-by-step cable selection process. 

1. System characteristics: 

a. Ac or dc. 

b. Grounding method (i.e., by grounding conductor 
or diode-grounding circuit). 

c. Normal operating voltage between lines or con- 
ductors (line-to-line voltage). 

d. Number of pilot or ground-check conductors and 
type of ground-check monitor. 

e. Minimum ambient temperature of cable storage 
and installation. 

f. Description of cable-installation area (surface 
mine, borehole, trailing cable, etc.). 

g. Environment of use (ambient temperature, 
amount of moisture, amount of sunlight, etc.). 

h. Maximum and normal operating current. 

i. Time schedule. 

j. Delivery point. 

k. Future changes in the system. 



2. Cable characteristics: 

a. Cable length. 

b. Cable type, number of conductors, and flat or 
round configuration. 

c. Voltage rating. 

d. Type of conductor (copper or aluminum). 

e. Conductor size. 

f. Insulation type. 

g. Jacket type and color. 

h. Maximum outside diameter and tolerance. 

i. Method of conductor identification. 

j. Special markings (MSHA and P approval num- 
bers, dating, etc.). 

k. End attachments (couplers), type of attachment, 
location of installer, and method of installation. 

Many of the items in the system characteristics cate- 
gory are obviously designed to assist the purchaser in 
identifying a specific cable type. For example, the number 
of power conductors is determined when ac or dc is 
specified (la). The need for one or more grounding conduc- 
tors is noted when the grounding method (lb) is explained. 
Similarly, the normal operating voltage (lc) leads to the 
selection of a cable voltage rating that includes the oper- 
ating voltage and the requirements for shielding. If the 
ground-continuity monitor requires a ground-check con- 
ductor, this should also be noted (Id). Any additional 
monitoring or remote-control systems may also require 
pilot conductors. Because cable jackets can crack during 
installation after being stored outside in extremely cold 
weather, the ambient temperatures of storage and use (le) 
should be specified. The installation area, category (If), 
explains special requirements such as high-tensile- 
strength conductors or a flame-resistant jacket for a bore- 
hole cable. Special environmental considerations (lg) that 
may affect cable life, such as an excessive exposure to 
sunlight in a surface mine, should be noted. Delivery time 
schedule (li) and delivery location (lj) are obviously im- 
portant considerations to be included so that a cable 
manufacturer can give the proper service. Finally, if 
changes to the electrical system (Ik) are anticipated, they 
should be considered. Money can be saved by purchasing a 



195 



cable that will accommodate both the present and future 
systems rather than replacing a cable after a short oper- 
ating period. 

Cable Length 

The second section of specifications is concerned with 
the detailed description of a required cable. First the cable 
length (2a) must be specified. Many companies prefer to 
purchase a long length of cable, thereby receiving a price 
discount, and then cut the required lengths from this 
stock. For instance, high-voltage feeder cable is usually 
shipped to a shop where couplers are mounted onto the 
cable at 1,000-ft intervals before the cable, now in the 
desired lengths, is transported to the mine. However, other 
factors such as Government regulations and voltage drop 
must be considered. Table 8.6 gives relevant information 
for underground trailing cables longer than 500 ft, based 
on a 60°C-rated insulation (a table for 90 °C insulation is 
not presently available) (38). 



Table 8.6.— Specifications for trailing cables longer than 
500 ft 



Conductor 
size 



Maximum allow- 
able length, 
ft 



Normal ampacity 

at 60°C copper 

temperature 

(40°C ambient), A 



Resistance at 
60°C copper 
temperature, 



AWG: 








6 


550 


50 


0.512 


4 


600 


70 


.353 


3 


650 


80 


.302 


2 


700 


85 


.258 


1 


750 


110 


.220 


1/0 


800 


130 


.185 


2/0 


850 


150 


.157 


3/0 


900 


175 


.130 


4/0 


1,000 


200 


.116 


MCM: 








250 


1,000 


220 


.098 


300 


1,000 


240 


.082 


350 


1,000 


260 


.070 


400 


1,000 


280 


.061 


450 


1,000 


300 


.054 


500 


1,000 


320 


.050 



Most of the remaining cable specifications have been 
discussed earlier in the chapter. Conductor size selection, 
however, is a complex topic that requires detailed analysis. 

Conductor Selection 

The selection of the conductor size (2e) is dependent on 
many parameters, such as ampacity, cable heating, volt- 
age drop, length, breaking strength, weight, shielding, 
insulation, and conductor material; the cable application 
may place emphasis on specific parameters. The correct 
selection will allow the cable to carry current without 
overheating or physical damage, to withstand the rugged 
mine environment, and to limit the voltage drop between 
the power source and the machines. 

Ampacity 

The ampacity or normal continuous-current rating of 
a cable is the current-carrying ability of its power conduc- 



tors. It is dependent upon the ability of the cable assembly 
to dissipate heat without damaging the insulation. The 
ampacity rating is usually based on the maximum conduc- 
tor temperature rise, with the temperature limit chosen 
on the basis of the specified life expectancy of the cable 
insulation. The temperature class assigned to the material 
used for the conductor insulation describes the maximum 
allowable sustained conductor temperature in a specified 
ambient temperature. The popular temperature ratings 
are 75° and 90°C. Cable insulation with a 60°C rating can 
still be found, but this value is no longer used extensively 
in mining. An ambient temperature of 40 °C is used for all 
ratings. 

The heat generated in the cable is primarily caused by 
the I 2 R power loss from current flow through the power- 
conductor resistance. The dissipation of this heat is a 
function of (30). 

• The conductor diameter and the number of conduc- 
tors in the cable; 

• The thickness of the conductor insulation and the 
cable jacket; 

• The cable configuration and outside dimensions; 

• The heat-transfer properties of the cable compo- 
nents; and 

• The type of conductor and cable outer jacket, and 
the ambient temperature. 

A conductor size (cross-sectional area) within a specific 
insulation and cable configuration is given a current 
rating (its ampacity) through calculations using these 
parameters and the generated heat. 

Cable ampacities are now designated in the United 
States by the National Electrical Code (NEC) (2) or by the 
ICEA for cables manufactured according to its design 
specifications. Parts 18, 75, and 77, 30 CFR, basically 
allow compliance with either the NEC or ICEA ratings 
(38). However, allowable ampacities for insulated conduc- 
tors given in the NEC are broad in both scope and 
application, and the same current value can be specified 
for one, two, or three conductors in a raceway, cable, or 
buried directly in earth (2, table 310-16). The broad 
applicability of the NEC standards implies that a safety 
factor must be built into its ratings, and comparison shows 
that the NEC ampacities are approximately 25% higher 
than the ICEA ratings. While the NEC values are fine 
within the scope and objectives of that code, ICEA values 
are preferred for engineered systems. Tables 8.7 and 8.8 
give the ICEA ampacities for the 90°C-rated cables pre- 
ferred for mining. Table 8.6 includes ampacities for 60°C- 
rated cables as specified in 30 CFR 18, and these are 
similar to the NEC values. 

The ampacity of a particular cable assumes that all 
splices, joints, and terminations in the cable are adequate 
in design and able to operate without restricting the 
loading on the cable. Considering the large number of 
splices made in mining cables, this assumption is a very 
important criterion for the cable rating. 

The ambient air temperature for the ampacities given 
in tables 8.7 and 8.8 is 40°C. If the maximum ambient 
temperature is different from that specified, the ampacity 
correction factors shown in table 8.9 should be applied 
(30). 



196 



Table 8.7.— Ampacities 1 for portable power cables, amperes per conductor 



Single conductor 

Conductor 0-2,000 V 2.001- 8,001- 15,001- 

size unshielded 8,000 V z 15,000 V 2 25,000 V 2 

shielded shielded shielded 

AWG: 

8 83 — — — 

6 109 112 — — 

4 145 148 — — 

3 167 171 — — 

2 192 195 195 — 

1 223 225 225 222 

1/0 258 260 259 255 

2/0 298 299 298 293 

3/0 345 345 343 337 

4/0 400 400 397 389 

MCM: 

250 445 444 440 430 

300 500 496 491 480 

350 552 549 543 529 

400 600 596 590 572 

450 650 640 633 615 

500 695 688 678 659 

550 737 732 — — 

600 780 779 — — 

650 820 817 — — 

700 855 845 — — 

750 898 889 — — 

800 925 925 — — 

900 1,010 998 — — 

1,000 1,076 1,061 — — 

1 Based on a copper conductor temperature of 90°C and an 

2 These ampacities are based on single isolated cable in air 

NOTE. — Dash indicates cable is not made. 



2-conductor, 
round and 

flat, 
0-2,000 V 



3-conductor, 
round and 

flat, 
0-5,000 V 
unshielded 



3-conductor round 



8,001 15,001- conductor, conductor, conductor, 
8,000 V 15,000 V 25,000 V -2,000 V 0-2,000 V 0-2,000 V 
shielded shielded shielded 



72 
95 
127 
145 
167 
191 
217 
250 
286 
328 

363 
400 
436 
470 
497 
524 



59 
79 
104 
120 
138 
161 
186 
215 
249 
287 

320 
357 
394 
430 
460 
487 



93 
122 
140 
159 
184 
211 
243 
279 
321 

355 
398 
435 
470 
503 
536 



164 
187 
215 
246 
283 
325 

359 



178 
191 
218 
249 
286 
327 

360 



54 
72 
93 
106 
122 
143 
165 
192 
221 
255 

280 
310 
335 
356 
377 
395 



50 

68 

88 

100 

116 

136 



48 
64 
83 
95 
110 
129 



ambient air temperature of 40°C. 
operated with open-circuited shield. 



Table 8.8.— Ampacities 1 for three-conductor mine power 
cables 



Conductor size 


2,001 to 8,000 V 


8,001 to 15,000 V 


Copper 


Aluminum 


Coppe 


Aluminum 


Copper 


Aluminum 


AWG 


6 





93 










— 


4 


2 


122 




124 


125 


128 


2 


1/0 


159 




165 


164 


168 


1 


2/0 


184 




189 


187 


192 


1/0 


3/0 


211 




218 


215 


221 


2/0 


4/0 


243 




251 


246 


254 


3/0 


250 


279 




278 


283 


281 


4/0 


350 


321 




342 


325 


344 


MCM 


250 


400 


355 




360 


359 


367 


300 


450 


398 




395 


401 


393 


350 


500 


435 




425 


438 


424 


400 


— 


470 




— 


473 


— 


450 


— 


502 




— 


504 


— 


500 


— 


536 




— 


536 


— 



1 Based on ICEA values with an ambient temperature of 40°C and a 
conductor temperature of 90°C [taken from "Power Cable Ampacities" (20), 
v. 1 for copper conductors and v. 2 for aluminum conductors]. 

NOTE.— Dash indicates cable is not made. 



Table 8.9.— Correction factors for ampacities at various 
ambient temperatures. 



Ambient 
temperature, °C 

10 

20 

30 



Correction 
factor 

1.26 
1.18 
1.10 



Ambient 
temperature, °C 



40. 
50. 



Correction 
factor 

1.00 
.90 



Cable Heating on Reels 

A cable that is used in a confined space can become 
overheated with continuous-current flow at the ampacity 
rating. Perhaps the best example is a cable bound on a 
reel, either for storage purposes or to increase mining 
machine mobility. Investigations were conducted as early 
as 1931 to identify factors responsible for overheating of 
rubber-jacketed cables, with emphasis on increased tem- 
peratures occurring in reeled cables (16). A cable manu- 
facturer manual published in 1940 was the first to contain 
a table of derating factors related to the number of layers 
wound on a reel to reduce the current-carrying capacity of 
the cable (23). These factors were included in ICEA spec- 
ifications for 60°C-rated cables in 1946 and have remained 
a standard since that time. Table 8.10 presents the ICEA 
values presently required by Federal regulations for all 
cable insulations (38). 

Research has been conducted since the publication of 
the ICEA derating factors to determine their applicability 
to the mining industry. McNiff and Shepherd (23-24) 
worked with cyclic currents, comparable to those experi- 
enced by shuttle cars in service, and steady-state loading 
at various percentages of cable ampacity, with both ac and 
dc power. Derating factors for 60°C-rated cables abstracted 
from these results are presented in table 8.10. An impor- 
tant contribution of their work, which cannot be shown in 
the table, is identification of the dependence of cable 
derating factors on the maximum-limit temperature per- 
mitted: at this temperature is increased or reduced, the 
derating factor changes accordingly. This was later veri- 
fied by Woboditsch (41), and his values for a limit temper- 
ature of 60°C are also given in table 8.10. 



197 



Table 8.10.— Ampacity derating factors for 60°C-rated trailing 
cables operated on drums 

McNiff and Shepherd (23-24) 1 Woboditsch 

Number of layers ICEA ,. 1s z 

ac dc y*'i 

1 0.85 3 0.78 3 0.82 0.78 

2 65 .54 .50 .62 

3 45 .42 .41 .48 

4 35 .36 3 .35 ( 4 ) 

5 NA .32. ^34 ( 4 ) 

NA Not available. 

1 Data for a 2-conductor, No. 4 AWG, type G cable at a maximum 
temperature of 60°C. 

2 Data for a 3-conductor, type NTSCE cable at a maximum temperature of 
60°C. 

3 Values from extrapolated curves since data did not extend to this range. 

4 Cable not made. 

Cable ampacity must be derated if the cable is used in 
a confined space. In view of the findings on limit temper- 
ature change, the ICEA values are probably adequate for 
75°C-rated and 90°C-rated cables. It is significant that 
Australian mining companies have recently accepted the 
initial derating factors, but with qualification (9), as 
shown in table 8.11. The ICEA values are specified as 
pertaining only to round cable, while new values have 
been generated for flat cable (8). As flat cable usually 
occupies more volume on a reel than round cable, heat 
transfer for flat cables should be less, and the lower values 
appear reasonable. 

Table 8.11.— Australian specifications for ampacity derating 
factors for trailing cables operated on drums 





Number of layers 


Circular cable 


Flat cable 


1 




0.85 


0.68 


? 




.65 


.52 


3 




.45 


.36 


4 




.35 


.28 



Current Calculations 

Current and voltage regulation are the two major 
concerns in sizing a cable correctly for an intended appli- 
cation. The effective continuous current through the cable 
power conductors must be less than the cable ampacity, 
with correct derating factors applied. The voltage drop 
across the distribution and utilization systems must be 
such that voltage regulation is within the tolerances 
specified for the loads. For trailing cables serving ma- 
chines, current is often the determining factor, since these 
cables are always short enough for voltage regulation not 
to be a problem. Feeder and portable cables serving many 
loads, however, are often so long that voltage drop becomes 
a principal concern. Even though the cable size may be 
found adequate in terms of ampacity and voltage drop, 
other factors may enter into the conductor sizing, such as 
tensile load, weight, and available short-circuit current. 

There are three basic methods that can be used to find 
trailing-cable ampacity: a full-load current similar to that 
specified in the NEC, a 30-min effective current demand, 
and a load-factor approach. Regardless of the method used, 
the engineer should realize that the typical current re- 
quirements of mining machinery change continuously 
over time and may be described as unsteady in nature. The 
infinite variability of mining conditions makes it difficult 
to define current levels for any part of a given duty cycle 
with precision. 



Calculation of cable ampacity requirements based on 
a 30-min effective current demand recognizes this vari- 
ability and also that cable heating varies as the square of 
current. Here, line current measurements are taken from 
the machine, and an effective or rms value is found by 
weighting current with 



El^t 
Et 



(8.2) 



where I effe c t ive = weighted current through cable, A, 

I = current level for specific increment of 
time, A, 
and t = time increment for current level I, s. 

This method does account for the transient heating and 
cooling of the cable, which should be considered for match- 
ing the loading conditions found in mining with the 
specific limit temperature for the cable; in other words, the 
ampacity. Through this method, representative machines 
in typical mining conditions can be measured and a 
catalog of effective currents can be assembled for ampacity 
selection. However, actual measurements are not always 
possible, and the next two methods do not require them. 

The full-load current approach is detailed by MSHA (39) 
and essentially follows the NEC requirements in sections 
430-22, 430-23, and 430-24. Here the ampacity of a cable 
supplying a single motor must be not less than 125% of the 
motor full-load current rating. When two or more motors are 
supplied through one cable, the ampacity must be at least 
equal to the sum of the full-load current ratings of all the 
motors plus 25% of the highest rated motor in the group. 
Provisions are allowed in this approach for adjusting the 
current requirements of any motor used for intermittent or 
periodic duty, and for the 60-min-rated motors normally 
found in mining (36); that is, the ampacity may be reduced by 
10% or 5%, respectively. 

The third method uses the machine load factor and 
applies the average power formula (32-33). For ac machines, 



I = 



HLF) 
V3 Vtj (pf) ' 



and for dc equipment, 



I = 



P(LF) 



(8.3) 



(8.4) 



where I = machine line current, A, 

P = (746) (hp) = rated average power of machine, 

W, 
hp = rated machine horsepower, 

_ _ actual average power consumed , . 

Lb = t— j = machine 

rated average power 

load factor, 

V = line-to-line machine voltage, V, 

pf = machine power factor, 

and 7j = machine efficiency. 

The formulas may be used for single motors or machines 
containing a complex of motors. Obviously, the load factor, 
power factor, and efficiency of a machine must be known in 
order to apply this metbod. With knowledge of typical 
operating conditions, these can be estimated. Values for 



198 



many underground coal mining machines have been re- 
searched and may be found in references 28, 32, and 33. A 
summary of these and values extrapolated from represen- 
tative underground mining conditions is given in table 
8.12; 100% efficiency should be assumed when applying 
these values to the formulas. However, caution should be 
taken when using these parameters as they are only 
representative. If precise currents are necessary, power 
measurements should be taken to obtain load factors and 
power factors, and manufacturer specifications consulted 
for efficiencies. The formulas can also be employed directly 
for full-load current calculations by assuming that pf = 
0.85, LF = 1 and 77 = 1 for ac induction machines, and pf 
= 1 and LF = 1 for dc motors. 

Table 8.12.— Some estimated power factors and load factors 

for various underground coal mining equipment in good 

operating conditions 

Machine Power factor 1 Load factor 2 

Battery chargers 1.0 0.8 

Belt drives .8 .7 

Belt feeder .8 .7 

Belt feeder breaker .7 .6 

Continuous miners 3 .6 .5 

Cutting machines 3 .7 .6 

Drilling machines .8 .7 

Lighting 1.0 1.0 

Loading machines 3 .7 .6 

Longwall shearing machines .8 .7 

Roof bolters .6 .3 

Section fans .7 .6 

Shuttle cars .6 .4 

1 For ac equipment only. 

2 For ac or dc equipment. 

3 Values are for cutting and/or loading only. Values for other machines are 
an average over a typical duty cycle. 



EXAMPLE 8. 1 

The difference between the last two methods can 
easily be seen through examples. First consider a 
150-hp ac continuous belt-conveyor drive motor 
rated at 550 V and operating in 20° C ambient 
temperature. Using the NEC currents (2, table 430- 
150) and applying 125% for the full-load current 
approach, the current used to size the cable would be 

I = (144X1.25) = 180 A. 

The ICEA ampacities of Nos. 2 and 1 AWG 3/C 
unshielded round cable from table 8.7, corrected by 
the factors in table 8.9, are (138X1.18) = 163 A and 
(161X1.18) = 190 A, respectively. Hence the No. 1 
AWG size would be indicated. Applying a load-factor 
calculation with table 8.12 data, 



I = 



(150X746X0.7) 
V3 (550X1X0.8) 



= 103 A. 



This relates that a No. 4 AWG 3/C unshielded round 
cable is adequate with a corrected ampacity of 
(104X1.18) = 123 A. The second method is probably 
more representative of actual conditions, since the 
NEC applies a 25% safety factor. 



EXAMPLE 8.2 

A cable size must be found for a 105-hp dc 
shuttle car. The machine is rated at 250 V, and it is 
assumed that the maximum ambient temperature is 
20° C, and an average of two layers of cable will 
remain on the reel. The load-factor approach will be 
used. 

From the information in table 8.12, a represen- 
tative load factor for shuttle cars is 0.4. Applying 
equation 8.4 and assuming 100% efficiency, 



I = 



(105X746X0-4) 
(250XD 



= 125 A. 



Ampacities for two-conductor cables from table 8.7 
corrected for a 20° C ambient temperature (table 
8.9) are 

• for No. 4 AWG, (127X1.18) = 150 A, 

• for No. 2 AWG, (167X1.18) = 197 A, 

• for No. 1 AWG, (191X1.18) = 225 A. 

This is a reeled application and these ampacities 
must be derated by the number of layers on the reel. 
Because of present Federal acceptance, the ICEA 
derating values from table 8.10 will be used. Thus 
for two layers, the ampacities must be reduced by 
0.65, or 

• for No. 4 AWG, (150X0.65) = 97 A, 

• for No. 2 AWG, (197X0.65) = 128 A, 

• for No. 1 AWG, (225X0.65) = 147A. 

Therefore, No. 4 AWG is too small, and No. 2 AWG 
would be selected. 

It can be noted that No 3 AWG was not included 
in the example. The reason is that this cable is not 
popular and is not readily available from manufac- 
turers. 



EXAMPLE 8.3 

Now consider a 550-Vac continuous miner that 
has five motors (two 50-hp gathering-head motors, 
two 175-hp cutter motors, and one 135-hp pump 
motor) for a total connected horsepower of 535 hp. 
Using the NEC currents (2, table 430-15), applying 
the intermittent-duty rating for the gathering head 
and cutter motors, and increasing the highest rated 
motor in the group by 25%, 

• 50 hp, I = (52X0.9) = 46.8, A 

• 135 hp, I = (133X1.0) = 133 A, 

• 175 hp, I = (168X0.9) = 151.2 A, 

• 175 hp, I = (168X0.9X1.25) = 189 A. 

Assuming the current phasor angles are such that a 
direct summation introduces only minor error, total 



199 



current for ampacity selection would be about 520 
A. Assuming the machine is operating in good 
mining condition, and using a load-factor calcula- 
tion with table 8.12 values, 



I = 



(535X746X0-6) 
V3 (550X1X0.6) 



= 419 A. 



Continuous miners of this size commonly use un- 
shielded 4/0 trailing cables with 90° C-rated insula- 
tion. If the ambient is 20° C, the ICEA ampacity 
from table 8.7 corrected with table 8.9 data is 
(287X1.18) = 339 A. This is considerably below the 
calculated values of 520 and 431 A. Actual visits to 
underground mines using continuous miners of the 
same size (535 hp) showed that the 4/0 cable jackets 
were not warm to the touch, implying cable- 
conductor temperatures well below the 90° C limit 
temperature (32). Furthermore, the load-factor cal- 
culation is based on data from machine cutting and 
loading, and since a continuous miner does not cut 
and load continuously, the current would be biased 
toward a worst case situation. Including the other 
machine operations (tramming, idle, etc.) would 
lower the load factor and the calculated current, 
probably below the ICEA ampacity. Regardless, the 
load-factor approach reflects this utilization envi- 
ronment more accurately than the NEC approach. It 
should be obvious that the effective current demand 
method would be more precise than either of these 
approaches. 



Intermittent Duty Ratings 

A major problem implied in the preceding example is 
that intermittent, fluctuating, or cyclic current through a 
cable has a different effect on cable heating than contin- 
uous loading. The full-load current or NEC approach for 
conductor sizing basically assumes continuous loading, 
but true continuous operation of most mining machinery 
would be a rare occurrence. Mining is inherently cyclic in 
nature. The Institute of Electrical and Electronics Engi- 
neers (IEEE) (17) does publish guidelines for rating elec- 
trical equipment under various operating conditions, du- 
rations, and time sequences of duty. Even though these 
terms have been used previously ifi this text, it is benefi- 
cial to define them here: 

• Continuous duty. Operation at a substantially con- 
stant load for an indefinitely long time. 

• Short-time duty. Operation at a substantially con- 
stant load for a short and definite specified time. 

• Intermittent duty. Operation for alternate intervals 
of load and no-load as definitely specified. 

• Varying duty. Operation where the amount of load 
and the length of time the load is applied are subject to 
considerable variation. 

In an endeavor to overcome the problem of mining 
duty cycles, the United Kingdom and Australian mining 
laws permit intermittent-duty ratings for mining trailing 
cables (9, 37). These ratings for several popular cable sizes 
are given in table 8.13. It can be noted that in both United 
Kingdom and Australian practice, the rating criteria are 



Table 8.13.— Intermittent-duty ratings for trailing cables 

Cable ,,I Xlm ? e Continuous Intermittent . „_ 

U.o. cable . . ncrease, 

size, . . , current current „. 

2 equivalent, »■ * .-a % 

mm * ^ AW _ rating, A rating, A 

UNITED KINGDOM 1 

16 5 85 90 6 

25 , 3 110 120 9 

35 2 131 145 11 

50 1/0 168 190 13 

70 2/0 205 235 15 

95 4/0 247 290 18 

AUSTRALIA 2 

21 4 70 95 36 

33 2 90 125 39 

1 Criteria: full-load current for 40 min, no-load current for 10-15 min, 1/2 
full-load current for 40 min, no-load current for 10-15 min; ambient at 25°C. 

2 Criteria: full-load current for 30 min, no-load current for 30 min. 



independent of the cable size. An attempt to match or 
classify the duty of mining machines with the well-defined 
IEEE categories, however, results in only one conclusion: 
the typical mining duty is equivalent to a varying-duty 
classification. Although mining sequences through given 
events regularly, distances constantly change; hence, 
equipment utilization changes. In such cases, the IEEE 
recommends the use of standard application methods to 
offset the problems of a nonconstant load, and suggests the 
use of load-factor and rms current calculations. These 
should be applicable to electrical equipment, such as 
cables, which are "sufficiently standardized both in per- 
formance and construction" (17). 

Voltage Calculations 

The major concern for voltage calculations is that 
adequate voltage must be at the machine terminals for 
proper starting and operation. As stated in chapter 6, the 
allowable voltage tolerance on all rotating machines is 
+ 10% for normal load conditions. Maintaining adequate 
voltage is one of the more difficult problems in mining, 
and is often the main constraint on mine expansion from a 
point of power delivery to the operation. 

As mentioned earlier, the voltage drop across trailing 
cables that have been properly selected by current calcu- 
lations is usually not a problem because of length con- 
straints in mining. This is especially true in underground 
coal mining, where the maximum length is restricted by 
the cable size used (as shown in table 8.6). One problem 
here, however, is that the maximum practical trailing- 
cable size that can be used is also constrained by the 
maximum weight that workers can physically handle. For 
three-conductor cables, this is considered to be 4/0 AWG, 
but use of 4/0 AWG can cause voltage-regulation restric- 
tions on high-horsepowered machinery. Trailing-cable volt- 
age drop may also be a concern in surface mines where 
utilization is at distribution voltage levels. 

Using the allowable voltage tolerance as a guide, good 
practice calls for limiting the maximum voltage drop 
under normal load conditions to not more than 10% of the 
nominal system voltage for each voltage level. For surface 
mines where machines operate at the distribution voltage, 
this would be equivalent to a maximum voltage drop from 
the substation secondary to the machines. In underground 
or surface mines containing power centers or a unit 
substation, this is not so apparent. Again, the maximum 
voltage drop must be restrained to 10%, but such a drop 



200 



can occur across the trailing cable alone. Consequently, 
the power-center or unit-substation primary must be 
maintained as close to its normal voltage rating as prac- 
tical, lb obtain this objective in practice can be a very 
difficult task, because power centers, for example, are 
usually at the extreme end of the distribution system. 
However, most mine power-center transformers are de- 
signed with two 2.5% taps above and below the rated 
primary voltage. Therefore, when voltage taps are avail- 
able, the maximum allowable voltage drop under normal 
load conditions in the distribution system (from the sub- 
station to the power centers or unit substations) is 10%. 

It is interesting to compare the 10% allowance with 
other electrical applications. For lighting, the NEC recom- 
mends 1.0% (2). Industries other than mining consider 2.0% 
as good-to-excellent regulation and 4.0% as satisfactory. 

For a thorough voltage-regulation study of a mine, the 
impedances of the source, the transformers, and all cables 
must be known. Tables 8.14 and 8.15 provide typical 
resistance and 60-Hz reactance values for popular mining 
cables (5); the missing parameters in these tables imply 
the cable is not popular or not considered suitable for 
mining usage. Manufacturer, power-equipment, and util- 
ity specifications must be consulted for other information. 



If cable sizes are not known, an assumption has to be made 
in order to carry out the calculations. Obviously, the loads 
on the power system must also be known. A circuit 
diagram must then be prepared and calculations per- 
formed to see if there will be adequate voltage levels at the 
loads. If calculated voltages are below those tolerated, 
system impedance must be reduced: the most convenient 
way is to increase cable sizes. Calculations are again 
performed to check for the desired result. In other words, 
the process is basically trial and error. It must be per- 
formed for normal load conditions; however, it is also 
recommended that calculations be made to ensure that 
critical motors can be started under worst case conditions. 
Even with a small system using the per-unit method, 
the computations can become so involved that accurate 
hand calculations are extremely time consuming or nearly 
impossible to obtain. Consequently, load-flow computer 
programs are the only answer; these are discussed further 
in chapter 10. However, there are some simple hand- 
calculation procedures that may be used for initial cable 
sizing, or for quick verification of voltage conditions in an 
existing system. These methods will be explored in the 
next example. 



Table 8.14.— Resistance and reactance of portable power cable 



SHD-GC, 
15 kV 



SHD-GC, 
25 kV 



R (ac), i O/Mft X L (60 HJ, 2 Q/Mft 

Conductor G-GC 

size 75-C 90°C gZgC SHD-GC, SHD-GC, SHD-GC, 

ft> ° W ° 2 kV 2 kV 5 kV 8 kV 

AWG: 

8 0.838 0.878 0.034 — — — 

7 .665 .696 .033 — — — 

6 .528 .552 .032 0.038 0.043 — 

5 .418 .438 .031 .036 .042 — 

4 .332 .347 .031 .035 .040 0.043 

3 .263 .275 .031 .034 .039 .042 

2 .209 .218 .029 .033 .038 .040 

1 .165 .173 3 .030 .033 .036 .039 

1/0 .128 .134 .029 .032 .035 .037 

2/0 .102 .107 .029 .031 .034 .036 

3/0 .081 .085 .028 .030 .033 .035 

4/0 .065 .068 .027 .029 .032 .034 

MCM: 

250 .055 .057 .028 .030 .031 .033 

300 .046 .048 .027 .029 .031 .032 

350 .039 .041 .027 .029 .030 .032 

400 .035 .036 .027 .028 .030 .031 

500 .028 .029 .026 .028 .029 .030 

600 .023 .024 .026 .027 .028 .030 

700 .020 .021 .026 .027 .028 .029 

800 .018 .019 .025 .026 .028 .029 

900 .016 .017 .025 .026 .027 .028 

1,000 .014 .015 .025 .026 .027 .028 

1 Criteria: a. Sizes 8 to 1 based on tinned copper 94.16% conductivity. 

b. Sizes 1/0 AWG and larger based on tinned copper 96.16% conductivity. 

c. Resistance increased by increments per ASTM B-172, Note 7 (3), to compensate for stranding factor. 

d. Skin effect calculated according to Arnold's Table, National Bureau of Standards Monograph 125 (29). 

e. Nominal cross-sectional areas. 

2 Criteria: a. Based on conductor dimensions given for class-H rope-lay conductors in table 2.5 of ICEA S-19-81 (21). 

b. Extruded-strand shield thickness, 0.015 in. 

c. Insulation thickness according to nominals given in Interim Standard 6 to ICEA S-68-516 (79). 

d. Diameter adder of 0.075 in to allow for semiconducting tape and metal-braid shield. 

3 Deviation from normal progression due to changes in insulation. 

NOTE.— Dash indicates cable is not made. 



0.044 


— 


.042 


0.046 


.040 


.044 


.039 


.043 


.038 


.041 


.036 


.040 


.036 


.039 


.035 


.038 


.034 


.037 


.033 


.036 


.032 


.035 


.032 


.034 


.031 


.033 


.030 


.033 


.030 


.032 


.030 


.032 



201 



Table 8.15.— Resistance and reactance of mine-power-feeder 
cable 



Conductor 
size 



R (ac), 1 Q/Mft, 
90°C 



X L (60 Hz), 2 Q/Mft 



MP-GC, 
5kV 



MP-GC, 
8kV 



MP-GC, 
15 kV 



AWG: 

6 

5 

4 

3 

2 

1 

1/0 

2/0 

3/0 

4/0 

MCM: 

250 

300 

350 

400 

500 

600 

700 

800 

900 

1,000 

1 Criteria: a. 
b 
c 



0.510 
.404 
.321 
.254 
.201 
.160 
.127 
.101 
.080 
.063 

.054 
.045 
.039 
.034 
.027 
.023 
.020 
.017 
.016 
.014 



0.041 
.040 
.038 
.037 
.036 
.035 
.034 
.033 
.032 
.031 

.030 
.029 
.029 
.029 
.028 
.028 
.027 
.027 
.027 
.026 



0.044 
.042 
.041 
.039 
.038 
.037 
.035 
.034 
.033 
.032 

.031 
.031 
.039 
.030 
.029 
.029 
.028 
.028 
.027 
.027 



0.042 
.041 
.039 
.038 
.036 
.035 

.034 
.034 
.033 
.032 
.031 
.031 
.030 
.030 
.029 
.029 



Based on bare copper 100% conductivity. 
Nominal cross-sectional areas. 

Resistance increased by increments per ASTM B-8, Note 3, 
to compensate for stranding factor, 
d. Skin effect calculated according to Arnold's Table, National 
Bureau of Standards Monograph 125 (29). 
2 Criteria: a. Based on conductor dimensions given for class B concentric 
stranded conductors in table 2.2 of ICEA S-19-81 (21). 

b. Extruded strand shield thickness, 0.015 in. 

c. Insulation thickness according to nominals given in Interim 
Standard 5 to ICEA S-68-516 (79). 

d. Diameter adder of 0.033 in to allow for semiconducting tape 
and copper-tape shield. 

NOTE.— Dash indicates cable is not made. 



EXAMPLE 8.4 

Distribution cables for a segment of an under- 
ground coal mine must be sized. A sketch of the 
situation is provided in figure 8.14 where the loads are 
two continuous mining sections. Voltages given are 
line to line. In-mine measurements and analysis of 
identical section equipment working in similar condi- 
tions have shown an effective current demand of 58 A 
with 0.8 lagging power factor at the power-center 
primary, when the continuous miner is cutting and 
loading. Maximum ambient temperature is 20°C. In a 
detailed study, the substation transformer impedance 
must be included. For the sake of demonstration, 
however, the 7,200-V line-to-line voltage at the substa- 
tion secondary will be assumed constant. The recom- 
mendation for allowable voltage drop is 10% across the 
distribution system. As the impedances of the feeder 
and portable cables must be known to make the 
calculation, a good place to start is to estimate line 
currents and make an initial cable selection by am- 
pacity. From the given information, 



Ii = 



I 2 = 



53 A. 



I 3 is related to I x and I 2 but is not necessarily equal 
to their sum, because of the diversity of mining 



Substation 

5MVA 

7% reactance 

69 kV : 7.2 kV 



10,000-ft 
feeder 
cable 

/ 



1,000-ft portable cable 
It 



Bus representing 

double-breaker 

switchhouse 



( 



Power centers 

for continuous 

mining sections, 

J T y ZQQ-M 

primaries 



1,000-ft portable cable 



Figure 8.14.— Simplified one-line diagram for situation 
described in example 8.4. 

operations. Chapter 4 presented the concept of de- 
mand factor (DF) where using a value from 0.7 to 0.8 
is considered reasonable for mining sections: 0.8 
corresponding to two sections and 0.7 to four or more 
sections. Therefore, 



or 



I 3 = 



I 3 = DFCIi + I 2 ) (8.5) 

(0.8X53 + 53) = 84.8 A. 



A 7,200-V system requires the use of 8-kV shielded 
cables, and the corrected ampacity for No. 6 AWG 
from table 8.7 or 8.8 and table 8.9 is 

ampacity = (93X1.18) = 110 A. 

This means that on a current basis the size is ade- 
quate for all distribution cables. Considering the pref- 
erence of the coal mining industry for using only 
portable cables for flexibility, ground-check conductors 
for ground-continuity monitoring, and 90°C insula- 
tion, an SHD-GC cable is indicated. Table 8.14 can be 
consulted for its impedance. It can be seen in the table 
that No. 4 AWG is the smallest 8-kV SHD-GC porta- 
ble cable readily available. Hence, a No. 4 AWG will be 
tried. Its impedance per 1,000 ft is 

Z cable = 0.347 + J0.043 Q 
= 0.35 \TA°_ fl. 

Referring to figure 8.14, the voltage drop across the 
distribution line conductors to either power center is 
(taking the power-center voltage as the reference 
phasor): 

V d = I 3 [10(Z cable )] + \ [l(Z cable )] 

V d = (84.8[ -36.9° X3.5|7.1°) 

+ (53|-36.9°X0.35|7.1°) 



or V d = 296.81 -29.8° + 18.6 | -29.8° 
= 3151-29.8° V 



As per-phase analysis is required to compare this 
drop with that allowed, the line-to-neutral voltage of 
the distribution system is used, or 



V ln = 



7,200 



= 4,160 V. 



The allowable voltage drop is 

V d allowable = 0.1(4,160) = 416 V. 



202 



Therefore, the 315-V drop using No. 4 AWG SHD- 
GC cables is tolerable. If the voltage drop were not 
acceptable, an increase in cable size would lower the 
impedance and the drop. 

This simple example had equal cable lengths to 
the loads, and currents operating at the same phase 
angle. It should be noted that typical mining sys- 
tems have many more loads, varying cable length, 
varying load power factors, and so forth, and the 
complexity of hand calculations will increase sub- 
stantially. Per-unit techniques are a tremendous 
help, but computer analysis is a much more efficient 
way to solve such problems. Nonetheless, the tech- 
niques shown here are useful for partial sizing or 
spot-checking distribution cables. 



strength because copper begins to elongate at that point. 
Federal regulations acknowledge the problem of exceeding 
the cable mechanical strength and mandate a minimum 
trailing-cable size for underground coal mine face equip- 
ment: No. 4 AWG for two-conductor dc cables and No. 6 
AWG for three-conductor ac cables (38). 

Short-Circuit Currents 

The emergency-overload currents that copper conduc- 
tors can withstand without serious insulation damage are 
shown in the graph in figure 8.15 (5). If the anticipated 
short-circuit currents are greater than those shown in the 
graph for the initial selection of conductor size, a larger 
conductor or a better grade of insulation should be chosen. 
Chapter 10 covers the calculation methods. 



Cable Mechanical Strength 



CABLE INSTALLATION AND HANDLING 



The tensile load on the cable should be determined 
from measurements in the mine, bearing in mind the 
problems discussed at the beginning of this chapter. The 
power-conductor breaking-strength data in table 8.16 
should then be consulted to assure that the conductor size 
is large enough to carry the tensile load (5). Two things 
must be considered when using this table. First, ground- 
ing and ground-check conductors should not support any of 
the tensile load, so the overall cable breaking strength 
should include only the sum of the power-conductor val- 
ues. Second, the working tension, especially in reeling 
applications, should not exceed 10% of the breaking 

Table 8.16.— Solid-wire breaking strength 

Conductor Hard— Medium— Soft- 

size, 65,000 psi 55,000 psi 40,000 psi 

AW G lb kg lb kg It) kg 

4/0 8,143 3,693.6 6,980 3,166.1 5,983 2,713.8 

3/0 6,720 3,048.1 5,666 2,570.1 4,744 2,151.8 

2/0 5,519 2,503.4 4,599 2,086.1 3,763 1,706.9 

1/0 4,518 2,049.3 3,731 1,692.4 2,985 1,354.0 

1 3,688 1,672.8 3,024 1,371.7 2,432 1,103.1 

2 3,002 1,361.7 2,450 1,111.3 1,928 874.5 

3 2,439 1,106.3 1,984 899.9 1,529 693.5 

4 1,970 893.6 1,584 718.5 1,213 550.0 

5 1,590 721.2 1,265 573.8 961.5 436.1 

6 1,280 580.6 1,010 458.1 762.6 345.9 

7 1,030 467.2 806.7 365.9 605.1 274.5 

8 826.1 374.7 644.0 292.1 479.8 217.6 

9 660.9 299.8 513.9 233.1 380.3 172.5 

10 529.3 240.1 410.5 186.2 314.0 142.4 

11 423 191.9 327 148.3 249 112.9 

12 337 152.9 262 118.8 197 89.4 

13 268 121.6 209 94.8 157 71.2 

14 214 97.1 167 75.7 124 56.2 

15 170 77.1 133 60.3 98.6 44.7 

16 135 61.2 106 48.1 78.0 35.4 

17 108 49.0 84.9 38.5 62.1 28.2 

18 85.5 38.8 67.6 30.7 49.1 22.3 

19 68.0 30.8 54.0 24.5 39.0 17.7 

20 54.2 24.6 43.2 19.6 31.0 14.1 

21 43.2 19.6 34.4 15.6 24.6 11.2 

22 34.1 15.5 27.3 12.4 19.4 8.80 

23 27.3 12.4 21.9 9.93 15.4 6.99 

24 21.7 9.84 17.5 7.94 12.7 5.76 

25 17.3 7.85 13.9 6.30 10.1 4.58 

26 13.7 6.21 11.1 5.03 7.94 3.60 

27 10.9 4.94 8.87 4.02 6.33 2.87 

28 8.64 3.92 7.02 3.18 4.99 2.26 

29 6.97 3.16 5.68 2.58 4.01 1.82 

30 5.47 2.48 4.48 2.03 3.14 1.42 



Cables must be installed and handled correctly in 
order to minimize damage from tension, bending, twist- 
ing, physical wear, cold, heat, and chemical reaction. 
Cable maintenance costs can be reduced, cable life im- 
proved, and safety enhanced by proper installation and 
handling. In other words, the considerable amount of 



100 
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Conductor ' copper 
Insulation • crosslinked 

polyethylene and 

ethylene 

Curves based on formula : 

[ff » =0-0297 *[$&] 
I = Short-circuit current, A 
A = Conductor area, cmil 
t - Time of short circuit, s 
T| = Maximum operating 
temperature (90*C) 

T 2 = Maximum short-circuit 
temperature (250°C) 
































































































































































































































































































































_ 










1 

























2 1 2/0 4/0 

I/O 3/0 

CONDUCTOR SIZE 



500 



1,000 



Figure 8.15.— Allowable short-circuit currents for insulated 
copper conductors. 



203 



engineering expertise expended in the design, manufac- 
ture, and selection processes can be wasted if the cable is 
not utilized properly at the mine. 

Borehole Cables 



should be utilized to prevent the clamps from loosening. A 
useful formula for determining the cable-clamp spacing is 



S = 



9DL 
W 



(8.7) 



The mining or electrical engineer may not have to 
plan and supervise the installation of a borehole cable 
frequently; however, since this may be the main power- 
supply cable for the entire or a large part of an under- 
ground mine, safety and production are highly dependent 
on use of the correct techniques. The term borehole cable 
comes from the common practice of installing a cable in 
the vertical borehole that has been drilled into an under- 
ground mine for the purpose of power entry. However, it 
applies to any cable that is vertically suspended into a 
mine, regardless of the opening in which it is placed. The 
typical location other than the borehole is a shaft. 

Considerable tension is imposed on borehole cables, 
depending on the weight of the cable and the depth of the 
mine. Proper conductor selection, installation procedure, 
and suspension method are necessary to assure that the 
cable provides trouble-free service for the life of the mine. 
Shaft cables are also subject to damage from moving skips 
and spillage. An extremely wet environment is often 
encountered, which may cause corrosion and icing prob- 
lems. In addition, safety precautions must be taken to 
keep the cable from breaking loose and falling into the 
opening during installation. If the power conductors have 
enough strength to support the weight of the cable during 
and after installation, messenger wires (wire ropes) with 
cable-gripping clamps are not necessary. Otherwise, a 
messenger-wire suspension method or a metallic-armored 
cable must be employed. 

If an unarmored cable such as an MP-GC is used, the 
following formula can be used to calculate the safety factor 
for the tensile strength: 



SF = 



0.80 AT 
W 



(8.6) 



where, A = total area of power conductors, in 2 , 

T = tensile strength of conductors, psi (24,000 psi 
for soft drawn copper and 40,000 for medium- 
hard drawn copper), 

and W = weight of length of cable to be suspended, lb. 

If the safety factor is greater than 7, an end suspension as 
shown in figure 8.16 may be used without messengers (18). 
Equal tensioning of the conductors is imperative. 

If messengers are needed, the wire ropes must be 
made from a corrosion-resistant material such as stainless 
steel. The typical system uses clamps or wire-type cable 
grips at specified intervals to secure the cable to individ- 
ual messengers. The cross-sectional area and tensile 
strength of each messenger must be such that it can 
support the total weight of itself, the clamp, and at least a 
cable portion. Proven and tested clamps of the best quality 
should be used, or they will become the weak link in the 
installation. The high gripping force necessary on the 
cable jacket should be spread over a large area so the 
jacket is not damaged by pinching. The clamps are often 
vulcanized to the jacket to prevent this. In addition, a 
jacketing material that is not subject to cold-flowing 



where S = distance between clamps, ft, 

D = cable diameter, in, 

L = clamp length, in, 
and W = weight of cable, lb/ft. 

Generally, clamp spacing is greater than 25 ft and should 
not be more than 100 ft. 

An armored borehole cable is used where depth or 
location precludes the use of messengers. The armor 
usually consists of a serving of steel or aluminum alloy 
wire typically placed over the cable jacket. If this type of 
cable is chosen, the armor carries the tensile load, and the 
tension safety factor can be determined by 



SF = 



0.60 (BS) 
W 



(8.8) 



where SF = safety factor, 

BS = breaking strength of each wire in armor 
multiplied by number of wires, lb, 
and W = weight of cable length to be supported, lb. 

The minimum safety factor for armored cable is 5. Ar- 
mored cable may be necessary in shaft installations as 



Drip loop in 
power conductor 




Insulated grounding conductor 
to grounding resistor 



Bond to casing 



Figure 8.16.— Representative end-suspension termination 
for borehole cable. 



204 



protection against jacket damage from skips, cages, and 
spillage. 

Cables can be installed either by raising or lowering. 
Messenger-supported cables are usually lowered into posi- 
tion as each messenger must be clamped at the top. 
Raising is often preferred for self-supported cables because 
of the need to have a brake on the surface as well as a 
pulling force at the bottom when a cable is lowered. In 
either case, the location should be straight and free of 
obstacles. It is also important to locate the cable in an area 
protected from any ground movement that may result from 
the mining operation. When a structure is used at the top 
to support the cable weight, it should not only be strong 
enough but also be placed on a substantial concrete base. 
Any sheave wheels utilized during the installation should 
be larger then the minimum bending radius specified. 
Rollers should be used to prevent jacket damage when the 
cable is dragged on rough surfaces and to minimize the 
pulling force by reducing friction. Crews working at the 
top and bottom should have a good communication system, 
and the personnel working at the bottom should be ade- 
quately protected from injury should the cable break loose 
and fall. 

Feeder Cable Installation 

The power-feeder cable must be located in an area 
that is protected from damage by mobile equipment. In 
underground mines, it is supported from the roof in 
regularly inspected fresh-air courses and haulageways on 
properly spaced insulated hangers, which may be sup- 
ported by a messenger wire. Messenger supports are 
usually installed at 20-ft intervals, and cable support clips 
are placed on a 3/8-in messenger wire at 4-ft spacings as 
shown in figure 8.17 (31). The recommended static load per 
clip is 100 lb. A 1-1/4-in-diameter hole is drilled in the roof 
to place a 6-in-long expansion shell bolt for the messenger- 
wire support. The cable must not come into contact with 
any combustible material. In underground coal mines, the 
cable must be guarded in any location where miners 
regularly work or pass under it, unless it is 6-1/2 ft above 
the floor or rail. Extra lengths of cable should be stored in 
large figure 8 configurations in a well-ventilated area. The 
bending radii recommended by ICEA, shown in tables 8.17 
and 8.18, should be observed for both mine power-feeder 
and portable cables when they are being installed (19-21). 
During installation, care must be taken not to twist the 
cable; that is, the reel should be turned so the cable is 
unrolled rather than pulled from the end of the reel. 
Finally, damage can be averted if a cable that has been 
stored on the surface during the winter is brought into the 
mine or a workshop to warm before being flexed. 

Recommended Handling Practices 

After cables are installed, proper cable-handling prac- 
tices can increase personnel safety and cable life. Other- 
wise, damage can easily occur, especially to trailing ca- 
bles, such as machine runovers, cutting by sharp edges of 
machine frames and stress clamps, and abrasion from 
sharp rocks and mine openings. Research has produced 
numerous recommendations to minimize this damage (10, 
14). These are presented in the following section and are 
divided between those directly applicable to underground 
mining and those for surface mining. It should be noted, 
however, that some recommendations apply to all mines. 



Table 8.17.— Recommended minimum bending radius, 

unshielded or unarmored cables, as a multiple of cable 

diameter 



Conductor insulation 
thickness, mils 


1.0-in diam 
and less 


1.001- to 
2.000-in diam 


2.001 -in diam 
and over 


155 and less 


4 

5 

NA 


5 
6 

7 


6 


170 to 310 


7 


325 and over 


8 



NA Not available. 

NOTE. — These limits do not apply to bending around curved surfaces in 
tension during installation. Larger bends are required for such installations. 



Table 8.18.— Recommended minimum bending radius, 
shielded and armored cables, as a multiple of cable diameter 

Cable type Minimum bending radius 

Armored: 
Flat tape and wire ... 12 times the overall diameter. 
Interlocked 7 times the overall diameter, except for tape- 
shielded cables and where a larger radius is 
specified for unshielded cables. 
Shielded: 

Tape 12 times the overall diameter. 

Wire Same as for portable cables unless the cable is 

flat-tape or wire armored. 

Portable 6 times the overall diameter for round cables or 

the minor dimension for flat cables for insulations 
rated at less than 5,001 V. The minimum is 8 
times the diameter for cables rated over 5,001 V. 

NOTE.— These limits do not apply to bending around curved surfaces in 
tension during installation. Larger bends are required for such installations. 



Cables in Underground Mines 

For reeled-cable applications, such as on shuttle cars, 
the cables must be anchored separately from the power 
equipment serving as the power source. The cable anchor 
points should be constructed so as to prevent personnel 
injury should the tie point pull out of position. When more 
than one reeled cable is at the tie point location, separate 
anchor points should be used for each cable. This will 
ensure that a cable will not whip dangerously should one 
of the anchor points fail. This precaution will also prevent 
subsequent cable damage. 

A shock absorber should be used between the reeled 
cable and the anchor point to reduce instantaneous cable 
tensions (jerking). The use of a rubber-tire shock absorber 
is adequate, provided that a cable clamp is employed 
rather than tying the cable to the tire. However, other 
types of shock absorbers may be more effective. Hydraulic 
pressure for the machine reel should be checked periodi- 
cally and set to manufacturer specifications to minimize 
instantaneous cable tensions. 

Backspooling is the process of moving a reeled-cable 
vehicle in a direction opposite from that for which it was 
primarily designed, for example, where a shuttle car dump 
point is beyond the tie point in a direction opposite to 
(outby) the mining face (inby). Research has found that the 
highest cable tensions occur during backspooling, result- 
ing from the sudden change in reel rotation as the shuttle 
car passes the tie point (14). Backspooling should be 
avoided, but if it is necessary, the cable anchor point 
should be located as far away from the travel entry as 
practical. This allows more time for the cable reel to 
change the rotation direction, and thus, cable tension will 
be less. 



205 




maximum 
spacing 




KEY 



A 


Feeder cable 


E 


Bulldog clamp 


B 


Dead-end hook 


F 


Expansion shell 


C 


Turnbuckle 


G 


Messenger wire 


D 


Cable clip 


H 


Sister hook 



Figure 8.17.— Messenger wire supports for mine power-feeder cable. 



Minimizing the number of cable friction points be- 
tween the tie point and the face will ensure the most 
effective use of a cable shock absorber located at the tie 
point. Friction points prevent the tensions from being 
transferred back to the tie point. When slack cable is 
reeled in, every precaution should be taken to minimize 
reel momentum to prevent jerking the cable when the 
slack cable supply is exhausted. Reeling in slack cable 
slowly and cautiously will help minimize the possibility of 
whipping the cable. Maintaining a smooth mine bottom, 
especially in the vicinity of the tie point, will help mini- 
mize instantaneously high cable tensions resulting from 
the shuttle car's bouncing over an uneven mine bottom. 
Minimizing shuttle car speed when rounding pillar cor- 
ners and passing the tie point will help prevent fast 
changes in reel momentum. Consequently, instantaneous 
cable tensions will be less severe. 

Minimizing the amount of excess cable stored on a 
reel will prevent heat buildup in the cable. Cable abrasion 
on the shuttle car can be reduced by assuring that all 
contact points are smooth and rounded. If possible, install 
rollers or sheave wheels at contact points between cables 
and shuttle cars to reduce abrasion and cable flexure. 
Avoid severely bending and twisting the cable at the tie 
point and elsewhere. A clamp should be used to limit cable 
bending at the tie point to 90°. Cable twisting between the 
machine and the anchor point can also be minimized by 
locating the tie point a maximum distance away from the 
machine travel entry. If possible, locate repairs to the 



shuttle car cable outby the tie point, where cable stresses 
are less severe. 

Recommendations for drag-cable installations are not 
as extensive as those for reeled cables but are just as 
important. First, the length of drag cable that is pulled 
should be minimized in order to reduce tension. Pillar 
corner edges should be rounded to prevent cutting or 
tearing of the cables. Precautions should be taken not to 
pull the cable over jagged rocks, timber, or other sharp 
objects that might damage the cable. 

There are some general practices that should be 
followed for handling all cables. Insulated gloves should 
always be worn, particularly when cables are energized. 
All cables should be stored in a warm environment during 
cold winter months. If storage facilities are limited, cold 
cables should be placed in a warm location for at least 24 
h prior to use. Small-gauge uninsulated wire must not be 
used to suspend cables from the roof, as it has a tendency 
to cut the cable jacket. All cable routes should be located 
in entries where they are safe from runovers. All cables 
should be checked periodically for damaged areas and 
electrical deterioration. Cables should be prevented from 
coming into contact with various oils, greases, or other 
contaminants that may deteriorate the cable jacket. When 
purchasing cables, make certain that they comply with all 
Federal and State regulations. In terms of jacket outer 
dimensions, this precaution will ensure effective use of 
packing glands and cable-laying devices. 



206 



Cables in Surface Mines 

Various types of equipment are available to assist 
with cable handling in surface mines, from insulated 
long-handled hooks to elaborate hydraulic reels and aerial 
crossover bridges. Despite this, considerable haulage, 
dragging, and hand-loading of cables onto sleds and trucks 
is still required in many surface mines. Superficial cable 
damage from abrasion is a common problem, as is cable 
crushing by mobile equipment. 

The following cable-handling recommendations for 
surface mines were detailed in a 1981 report to the IEEE 
(10). Systems should be developed for clearly marking 
cable lines along roadways and in pit areas. Suitable 
crossovers should be provided; in heavy traffic areas, these 
should be elevated. Sleds, skids, reels, and so on should be 
utilized rather than dragging the cable. Nylon rope or any 
device that can kink the cable should be avoided. Strain 
relief should be provided where cables are attached to 
equipment; rope or wire cable should not be used for this 
purpose. Insulated gloves are in poor condition at many 
minesites and provide inadequate protection for cable 
handling. In addition, personnel tend to place the cable 
across the body, negating any protection afforded by the 
gloves. Tools designed for cable handling should be clean 
and in a good state of repair. They should always have 
insulated handles. 

Conroy and Mertain (10) have made a very important 
statement about cable handling that is applicable to all 
mines: "A training program for all persons engaged in 
cable handling should be mandatory. This should cover 
both electrical precautions and procedures— particularly 
de-energization and lockout— and physical methods. Cable 
handling tools and devices should be made available to all 
concerned, and their use should be mandated. Mechanized 
cable handling equipment should be considered from both 
a safety and an economic viewpoint; and it may occur that 
an actual cost saving can be demonstrated for its use." 



CABLE FAILURES AND REPAIRS 

Most electrical cables used in mining are designed to 
have a minimum life expectancy of 20 yr, with a safety 
factor of about 2. The life expectancy is controlled prima- 
rily by the service life of the insulating jacketing materi- 
als, which, as noted earlier, are temperature related. 
Where specified operating temperatures are exceeded, 
deterioration of the insulating materials is accelerated 
and the useful service life is shortened accordingly. Tem- 
perature is the main factor in the deterioration of nonport- 
able cables that are fixed in place for extended periods of 
time, provided that proper installation practices are fol- 
lowed using good techniques. Portable cables, on the other 
hand, are frequently exposed to both excess generated 
temperature and mechanical abuse. As a result, portable 
cables can experience repeated failures at frequencies 
directly related to the proximity of the cable to the active 
mining area, the general mining conditions, and mainte- 
nance and cable-handling practices. For portable cables, 
the design life of 20 yr can easily deteriorate to 1 or 2 yr of 
actual in-service use. 

Cable deterioration due to overheating is a time- 
dependent function and can go unnoticed in routine min- 
ing operations. The main indication is that the cable 
becomes uncomfortably hot to the touch or, in more severe 
cases, produces smoke or steam in wet conditions. Excess 



cable on a reel, created, for instance, by not taking into 
account cable derating factors, is the most probable con- 
tributing cause of cable overheating. 

Mechanical wear can also be a time-dependent factor in 
cable failures, as, for example, repeated abrasion on a sheave 
support or spooling eye on a shuttle car. The most likely 
causes of failure, however, are those abuses associated with 
immediate or nearly immediate power interruptions. A 
prime example is the case where a shuttle car operator 
exceeds the length of the car umbilical, and the cable is 
tensioned to the point of failure. Similarly, a shuttle car 
might run over its own cable, pulling it apart or crushing the 
conductors and insulations. One machine running over the 
cable that powers another machine is also a common abuse 
that eventually, if not immediately, takes its toll. Obviously, 
special care and consideration are needed to adapt such a 
relatively vulnerable item as a power cable to the mining 
environment. Unfortunately, once a cable has been damaged 
to the point of requiring a repair, it becomes more vulnerable 
than ever, since it is almost impossible to restore its original 
performance characteristics. 



Cable Testing 

Although cables are often not tested routinely at a 
mine, there are instances where testing is recommended. 
Manufacturers test the components used in manufactur- 
ing and do carry out limited testing of completed cables as 
prescribed by ICEA standards (19-21). When couplers are 
added to cables at a cable repair shop, further testing can 
be done and the person responsible should ensure that 
these tests are performed effectively. The mine should 
require every cable removed from service to be tested 
before re-installation. If this were done, many costly 
in-service failures, production losses, and safety hazards 
could be prevented. 

Visual observation of cable condition is an important 
and simple task that can be carried out even when the 
cable is in service. It is important to require machine 
operators to walk to and from their equipment along the 
cable and visually examine the jacket for damaged areas. 
Outside diameter and hardness can also be determined on 
in-service cables. Any significant reduction in the overall 
diameter is an indication of excessive tension, while 
increased hardness results from excessive temperature or 
bending. 

More extensive evaluations can be made when the 
cable is out of service and the conductor ends are accessi- 
ble. Obviously, an ohmmeter can be used to test for broken 
conductors; however, more sophisticated equipment is nec- 
essary to locate an open circuit. Insulation damage may be 
detected by using a megohmmeter or a high-potential 
tester (hipot), each of which can give an indication of the 
ability of the insulation to withstand the operating volt- 
age without allowing excessive line-to-line or line- 
to-ground leakage currents. Portable megohmmeters and 
dc hipots can be used in the field, and ac hipots are 
sometimes available at cable repair shops. In order to test 
nonshielded cables completely, they must be surrounded 
by a grounding medium such as a water bath; otherwise, 
only the insulation directly between power conductors and 
between power and grounding conductors can be exam- 
ined. If a shield or armor is present, either can be used as 
a grounding medium for the test. 

Two basic types of insulation testing can be accom- 
plished with these methods: acceptance and maintenance. 



207 



For acceptance testing, the ICE A standard procedures and 
voltage levels should be followed (19-21). Maintenance 
testing requires lower voltage levels to avoid damaging the 
cable during the test. In both tests, the voltage level and 
duration of test should be adequate to ensure that the 
cable will perform safely in the intended service. As a 
general rule-of-thumb, maintenance test voltage is at 50% 
to 70% of the ICEA acceptance test values and should be at 
least as high as the cable rating. Insulation resistance 
values from megohmmeter testing and leakage currents at 
specified test voltages, obtained from dc hipot testing, can 
be used for preventive maintenance scheduling. If records 
are maintained, these tests can be used to indicate replace- 
ment schedules and prevent in-service breakdowns. 

Failure Location 

Failure location, often termed fault location, is an- 
other type of testing that is extremely important because 
of the susceptibility of mine cables to damage. It is less 
time consuming to repair or splice a cable in the mine 
than to replace it, and it is essential to have quick and 
accurate methods for locating cable failures in order to 
minimize the loss of production time. The Bureau of Mines 
has evaluated several methods, some of which follow (11). 

Some faults are low-resistance short circuits that can 
be found by visual inspection. Nonvisible short circuits 
can be blown by applying a high-energy power source to 
make them visible. However, this practice is not recom- 
mended within mines because of the potential safety 
hazards of fire and personnel injury. When there are faults 
in more than one place or when they cause low-resistance 
open circuits or high-resistance short circuits, they are 
extremely difficult to locate. 

A thumper or capacitance-discharge fault locator has 
been used successfully in surface mines and in cable repair 
shops; however, associated safety hazards restrict its use in 
underground mines. A capacitor is charged until a spark 
gap breaks down, sending a pulse along the cable. If the 
resistance is low enough, the pulse will discharge across a 
short and return. The pulse will not propagate across a 
high-resistance open circuit at the same intensity as it was 
transmitted, and an acoustic sensor can be used to locate 
the area where the signal caused by the pulse became 
diminished. 

The time-domain reflectometer (TDR) is another 
fairly successful method for locating failures. It works on 
the principle of a reflected pulse that either reinforces or 
reduces the original signal, depending on whether the 
discontinuity is an open or a short. The time of arrival of 
the echo is proportional to the distance to the failure, and 
the distance is then visually displayed on a meter. An 
accessory probe is necessary for exact failure location 
when a TDR is used, since the precise measurement of 
distance along a cable is difficult in a mine. A tone 
transmitter can be used in conjunction with an audio 
probe to locate the failure precisely. An infrared probe can 
also be used to locate faults where temperature increases 
are evident. Probes sensitive to 1° or 2° F are available; 
however, a current source must be attached to the cable 
end. 

Splicing 

Once a cable is damaged and made unsafe or inoper- 
able, the damage must be repaired so that the machine 
might be put back into service with the least delay In U.S. 



mines, repairs of this type can be made on the spot, 
whereas in some countries, such as the United Kingdom, 
the cable is replaced in its entirety and transported to a 
cable repair shop. The Code of Federal Regulations (38) 
states that "temporary splices in trailing cables or porta- 
ble cables shall be made in a workmanlike manner and 
shall be mechanically strong and well insulated." It fur- 
ther states that "when permanent splices in trailing 
cables are made, they should be: 

• Mechanically strong with adequate electrical con- 
ductivity, 

• Entirely insulated and sealed so as to exclude 
moisture, 

• Vulcanized or otherwise made with suitable mate- 
rials to provide good bonding to the outer jacket." 

By Federal regulations, only one temporary splice is 
permitted in any one cable at any given time, and this 
must be removed or repaired within 24 h. A permanent 
splice, as the name suggests, can remain in place indefi- 
nitely so long as it is safe and effective. The number of 
permanent splices in a cable is not limited, except by 
Pennsylvania law where no more than four permanent 
splices are permitted along with one temporary splice. In 
other words, a trailing cable may contain five splices but 
only for a maximum time of 24 h. 

By law, specially approved splice kits or materials 
must be used when making a permanent splice repair. 
These kits and materials are tested and approved by 
MSHA and given an approval number similar to the 
approval number for cables. As with cables, a P is added to 
the MSHA number to signify approval for use in Pennsyl- 
vania. Depending on their basic components and outer 
coverings, splice kits are generally classified as tape 
splices, cold-sleeve splices or heat-shrink splices. Varia- 
tions of these three types depend on the manufacturer. 

Tape splices use tape for the conductor insulation 
components as well as for the outer jacket replacement 
materials. In some cases, slit insulation tubes might be 
used with the insulating tape and in other designs, 
blanket-type wraps might be applied in combination with 
moisture sealing tapes to provide an overall covering prior 
to applying the final layers of tape that form the jacket 
replacement. 

Cold-sleeve splices include a variety of conductor 
insulation materials and usually consist of tapes that are 
used alone or in combination with slit-tube insulations. 
The main differences lie in the method of applying the 
outer jacket replacement. In all cases, a sleeve is slid onto 
the cable prior to making the conductor connections. In 
some designs, the splice area is built up using insulating 
tapes, and then a generous amount of adhesive is applied 
over the tape. The adhesive also serves as a lubricant and 
so the tubular covering must be moved into place imme- 
diately to cover the splice. This covering is designed to be 
slightly undersized so that it stretches as it is placed over 
the bulky taped area. The sleeve must be pushed into place 
or grasped at the end for pulling, otherwise additional 
drag is produced in grasping the sleeve, and it might be 
almost impossible to position it properly. 

Some of the cold-sleeve splice coverings are pre- 
stretched during the manufacturing packaging process. 
When it is time to place the sleeve, it is merely moved into 
place over the splice area and the restraining device is 
removed, thus allowing the sleeve to shrink (recover) onto 
the splice area, which has a smaller cross section. One 



208 



such device holds the sleeve in an expanded position by use 
of an inner plastic core, which is progressively collapsed 
along the length of the sleeve. In another design, the 
sleeve is held expanded by an adhesive bond between it 
and a rigid external concentric tube. A solvent is applied 
between the sleeve and the tube when it is time to allow 
the sleeve to recover onto the splice area. In still another 
design, the covering comes with both ends prerolled to- 
ward the middle in a toroidal fashion. The covering re- 
mains in this configuration until it is unrolled over the 
splice area. All of these special cold-splice coverings are 
designed to facilitate placement of the covering. The 
adhesive bonding and moisture sealing vary from manu- 
facturer to manufacturer, with the necessary components 
included in the kit. 

The heat-shrink splice coverings are also prestretched 
but in a different sense. The sleeves are made of special 
cross-linked polymers, which stretch readily when warm. 
In the manufacturing process, they are heated to 270°F 
and expanded radially to a given oversized dimension and 
then cooled. While at room temperature, they retain this 
oversized dimension, which easily accommodates place- 
ment over the spliced cable. When reheated to 250°F, the 
sleeve will shrink onto the splice area. A factory-applied 
thermal-melt adhesive on the inside surface of the sleeve 
softens with the applied heat and forms a moisture seal 
and adhesive bond between the sleeve and the original 
cable jacket. Similar smaller sleeves are used for conduc- 
tor insulation over the individual power conductors. 

The packaged splicing kits contain all the materials 
necessary to reinsulate and rejacket the splice area, to- 
gether with special illustrated instructions. Cleaning ma- 
terials, a cloth, and a can of solvent might also be 
included, together with an emery cloth (nonconducting) or 
scraper, which is used to prepare the surface of the cable 
jacketing for improving the adhesive bond. The connectors 
used to rejoin the power, grounding, and ground-check 
conductors may or may not be included in the kits, 
depending mainly upon customer specifications. 

The Bureau of Mines has sponsored research into the 
causes and prevention of splice failures, with emphasis on 
shuttle car cables (26, 34-35). Since this research has had 
a positive influence on the splice kits and insulation 
procedures used in the mining industry, a brief overview of 
the results is presented. 

Deenergizing Procedures 

In the interest of safety, it is essential to follow strict 
lockout procedures before cutting into any cable that has 
been put into use. Improper lockout prior to splicing cables 
has been a major source of electrocutions in the mining 



industry. The individual making the repair must go to the 
power-source end of the cable, disconnect the cable, and 
tag (danger off) and lock out the disconnecting device, 
which is usually a coupler. This step must never be left for 
someone else to do. 

Cable Preparation 

After the cable has been properly disconnected from 
the power source, the next step is to remove the damaged 
area and prepare the conductors for splicing. The prepara- 
tion procedures vary slightly depending on the types of 
connections used, and a representative procedure is pre- 
sented in figure 8.18. A guide or template is recommended 
for marking the cable for cutting and removing the insu- 
lation and jacketing. Such a guide is included in some kits 
but can be easily fabricated from light-colored material for 
repeated use. Use of a marking guide can help to standard- 
ize procedures and increase speed and accuracy. 

Once the cable pieces are properly marked, the next 
step is to remove the unwanted insulation. An effective 
method is presented in figure 8.19. The key here is to use 
a sharp knife and to take care not to cut all the way into 
the conductors. Nicking the conductor strands will mini- 
mize their performance. The conductor connections are 
usually staggered to help reduce bulkiness (fig. 8.20). The 
marking guide maintains good positioning, and before the 
insulation is actually cut into, the guide allows an imme- 
diate check point to determine that the power conductors 
are properly registered; that is, black to black, white to 
white, and so forth. 




Figure 8.18.— Splice layout using template for staggered 
connections. 




\WW N 




X^8fc ; 


v\\\\S 




\. \r >w* * • iyl 


l£v\ 




\ T^ «V^^ D— ^ 


'"'^Am 


x^V\\ 


^fSucJi 


■*^-m/.^T 










iH 




^^ 



^S^-. *'& 




n 


W§M 


\\ffi 


K^ 





Figure 8.19.— Effective method for removing unwanted insulation. 



209 





Modified 
crowsfoot 




Figure 8.20.— Staggering splice connections. 



Connectors 

A variety of connectors (fig. 8.21) and connector crimp- 
ing tools are available. It is generally recommended that 
lapped-joint connections be used where maximum tensile 
strength is desirable, as in shuttle car cables. Research 
has shown that the modified crowsfoot connection, when 
properly installed, can restore 80% to 100% of the original 
tensile strength for Nos. 6, 4, and 2 AWG conductors, the 
smaller conductors being easier to restore. The modified 
crowsfoot connection offers additional advantages of axial 
symmetry (no mechanical couple) and a small profile (an 
important consideration with multiconductor cables). 

The lap joints, being shorter than butt joints, are 
better for reeling applications since repeated flexure on a 
long connection might accelerate fatigue failures. The lap 
joints generally outperform butt connections in tensile 
strength, and Bureau-of-Mines-sponsored research has 
shown that restoring tensile strength is probably more 
important than restoring high flexibility to shuttle car 
cables. Either way, the lap connections are superior. 

A major consideration in obtaining high tensile 
strength is the use of the proper crimping tool for a given 
connector. Furthermore, tools that reduce or eliminate 
operator judgment tend to provide the best repeatability, 
since overcrimping as well as undercrimping can reduce 
tensile values. 

The lap connection has also been recommended for 
restoring the grounding conductors. In this case, it has been 
suggested that the connection be a little forgiving and allow 
the grounding conductors to slip slightly inside the connec- 
tor should the cable undergo excess tension. This would 
cause the power conductors to take all the tension and would 
perhaps prevent the grounding conductors from being ten- 
sioned, so that they would be the last to fail. Although this 
concept has not been verified, it may have some merit 
(assuming of course that a good electrical connection is 
maintained and that otherwise the grounding conductor 
might not extend sufficiently under tension). 

An important consideration in selecting and install- 
ing connectors in reeled cables is awareness of the connec- 
tor profile after installation. Bulky connectors with abrupt 
edges are more difficult to insulate effectively, simply 
because they tend to cut through the insulation materials 
with repeated cycling under normal operations. These 
connectors can also cause excess pressure and fatigue on 
adjacent grounding conductors, which are uninsulated 
and somewhat less protected from mechanical abuse. 

Although generally unsatisfactory for related applica- 
tions, the butt connection is effective for larger portable 



Parallel 




Slide connector into 
place and crimp 



Full crowsfoot 





Butt 
(where tensile 
strength is not 
most important 
consideration ) 



Figure 8.21.— Examples of popular connectors and connec- 
tions used in splices. 



cables such as those used for continuous miners, because it 
offers the least bulk. Here it does not need to withstand 
the repetitious flexing so often experienced by the smaller 
size cables. 

Reinsulating 

Because of the repeated bending stresses, reinsulat- 
ing procedures require special attention in portable ca- 
bles. The key is to provide a flexible joint and seal where 
the new insulation contacts the original cable insulation. 
As shown in figure 8.22, this is best accomplished using 
soft rubber tape that completely fills the volume and laps 
over the original insulating material. The lap is important 
since a tape fill that only butts to the insulation is almost 
sure to separate after very little flexing. Where it is 
desirable to use slit tubes as part of the reinsulating 
procedure, soft tape is recommended underneath and over 
the tubing. 

Soft rubber tape alone will not hold up under repeated 
cable flexing. Therefore it is further recommended that 
tougher vinyl tape be applied over the rubber tape. The 
vinyl tape accomplishes two objectives: it restrains the soft 
tape, thus preventing it from squeezing and extruding 
from its intended area, and it allows the reinsulated 
connections to slide relative to one another and the 



210 



Cable insulation 



Connector 




Vinyl electrical tape 



Figure 8.22. — Reinstating power conductors with soft rub- 
ber tape. 



grounding conductors with minimum wear. The vinyl tape 
can also be used to bind the multiple conductors together 
for maintaining positioning and limiting excess relative 
motion. A single-width wrap of tape near the middle of the 
splice area is generally sufficient. Care should be taken 
not to use too much vinyl tape over the splice area, since 
the final splice covering is generally intended to bond to 
the inner parts and the vinyl can in some cases make the 
subsequent adhesive bond less effective. 

In the case of heat-shrink splices, the conductor insu- 
lations are also made of heat-shrinkable tubing, and the 
tubes must be slipped onto the conductors before the 
connector is applied. When shrinking the tubes with a 
heat source, care must be taken to avoid overheating or 
rupturing the insulation on the sharp connector edge, and 
so forth. After heating, the installer should inspect the 
work to ensure that the adhesive has sealed the sleeve to 
the original insulation material. This is especially impor- 
tant for flat cables where the insulation cross sections are 
not always smoothly continuous. The heat-shrink insula- 
tion tubes provide a generous lap over the original insu- 
lation and are usually tough and resistant to rubbing wear 
inside the splice. 

Shielded cables require complete shield replacement 
over the conductor insulation. This process is similar for 
all cables but requires more care in high-voltage splices 
and will be covered later. 

Rejacketing 

The outer splice covering provides protection for the 
more delicate inner splice components and serves basi- 
cally the same purpose as the rugged cable jacketing. It is 
important that it be tough and flexible and at the same 
time maintain an acceptable bond to the original jacket- 
ing material. Of principal concern is a splice condition 
generally termed end lipping, the result of the splice- 
covering ends' pulling away from the cable jacket. When 
this occurs, contaminants such as fine solids and water 
can enter the splice and contribute to failure or an unsafe 
condition. The causes of end lipping are combinations of 
poor adhesive bonds, discontinuities and dissimilar mate- 
rials, or simply physical wear as a result of the normal 
mining process. The amount of end lipping will vary 
depending on the types of covering used and the conditions 
to which it is exposed. 

Various attempts have been made to provide splicing 
products that resist end lipping, with varying degrees of 
success. The general recommendation is to prevent occur- 
rence by making every effort to clean the cable surface 
where the adhesive bond is to be made. As a minimum, 
any soiled surfaces should be wiped with a suitable solvent 
and abraided with nonconductive emery material to reveal 



a fresh bonding surface. It should be noted that newer 
cable jackets can be more difficult to bond simply because 
waxes from the manufacturing process are often on the 
jacket surface. 

In general, the heat-shrink sleeves are good abrasion- 
resistant coverings. However, it should be noted that they 
are usually stiffer and sometimes require more attention 
to obtain a good and lasting bond to the cable jacket. 
Furthermore, a heat-shrink sleeve can take on a thermal 
set, for example, if it is allowed to cool in a curved position 
on a reel and then is later unreeled while still cool. The 
cold splices are generally quite flexible, but end lipping 
can result from bending and scuffing on various machine 
parts. Major cable and splice wear usually occurs during 
contact with the machine and its spooling mechanism 
when the relative motion is at a maximum. 

It is normal practice to tape down the ends of the 
splice coverings. This can help to reduce end lipping and 
can also prevent foreign matter from entering an already 
lipped end. Regular inspection and renewal of the end 
taping is a must, since abrasive wear and cutting on 
machine parts can quickly destroy even well-applied end 
tapes. The use of exposed soft rubber tapes is considered 
poor cable repair practice. The softer rubber tapes can 
provide good moisture seals but should be protected with 
an overcovering of tough vinyl tape. This vinyl tape will 
help contain the rubber tape, and the lower friction will 
give better wear characteristics. 

High- Voltage Cable Splices 

When splices are required on high-voltage cables in 
underground mines or in surface mines, problems are 
introduced by the presence of shields and semiconducting 
layers. The high voltage means that care must be taken to 
achieve an excellent splice, that is, one that closely ap- 
proximates the qualities of the original cable. 

The splicing procedure is basically the same as that 
just covered, but the cable insulation and jacket are 
usually tapered as shown in figure 8.23. Tapering is 
performed to improve the bond, increase the leakage path 
length, and lessen the chance of a direct vertical path to a 
ground plane. Extra care and skill are necessary as any 
damage to the insulation during splicing, such as a small 
cut, will cause more rapid dielectric failure at higher 
voltages. In the same context, a small protrusion such as a 
sharp edged connector or loose wire will be a more 
noticeable failure initiator as the voltage increases. 

The presence of semiconductive tape and braid or tape 
shielding in cables requires extra caution. The shielding 
system must be separated from the conductor insulation in 
such a way that residue on the insulation from the 
semiconducting tape is completely removed before the 
conductor insulation is reapplied. In addition, the wires 
from a braid shield must not protrude into the insulation. 
The shield must be replaced completely, and the grounding 
conductor must be placed in intimate contact with the 
shield. 

Splice Inspection 

A recommendation for improving splice performance 
is to inspect splices on a regular basis and use the 
information to institute new procedures or even new splice 
kit designs. An opportune time for doing this is before 
shipping an extensively damaged cable to a repair shop for 
vulcanized repairs. When the cable is idle and quite 



211 



Outer protective 
cover tapes 



Cable shielding 




Grounding lead 
if necessary 



Figure 8.23.— Typical taped splice in high-voltage shielded 
cable. 



accessible, perhaps stored in a supply yard, old splices can 
be cut out and scrutinized. Just the simple process of 
slitting the old splice lengthwise using a sharp linoleum 
knife can provide good information regarding insulation 
procedures, wear characteristics, effectiveness of bonds, 
and so forth. Electrical tests and tensile evaluations can be 
made on the insulations. Samplings of this type can readily 
provide extensive data on splices with varying amounts of 
in-service time. 

TROLLEY SYSTEMS 

The conductors that provide power for electric track 
haulage systems form a major part of the power- 
distribution system in many underground mines. The 
trolley system is a potential hazard for fires, ignition of 
methane, and shock since it utilizes uninsulated conduc- 
tors. The danger in underground coal mines is greater 
than that in surface mines because of limited space and 
the presence of methane, However, all mines that utilize 
trolley conductors can benefit from proper design, selec- 
tion, and installation of the system components. 

Several conductors are used in the trolley circuit: 
trolley wire, feeder cable, rail-bond cable and steel track 
rails. The trolley wire supplies power directly to a rail- 
mounted vehicle, such as a mine locomotive, through a 
collector called a shoe or harp. The trolley wire and 
collector connection can cause frequent severe arcing, 
which may damage either part and cause an obvious 
ignition hazard. Proper positioning of the trolley wire, 
particularly at curves and switches, correct holding force 
on the collector, and the required amount of lubrication 
are necessary to minimize arcing. 

A feeder cable supplies power to the trolley wire. 
Consequently, both must be sized properly to provide 
enough current-carrying capacity yet minimize heating 
and voltage drop. In addition, rectifiers must be positioned 
at adequate intervals to supply the proper voltage to the 
feeder. 

The current return path utilizes the steel rails, which 
must have adequate conductivity to minimize the total 
system resistance. Rails are laid in segments, and the 
connections between them can loosen or the rails could 
break; hence, rail-bond cable is installed to maintain 
continuity. Rail-bond cable is attached at each rail joint, 
and as a further precaution, between the two rails at 
specified intervals (cross bonds). 

Trolley Wire 

The trolley-wire conductor used in mines is hard- 
drawn copper, but brass is available for high-speed surface 



transportation. Round, grooved, figure 8 and figure 9 
(deep-section) wire shapes, shown in figure 8.24, are avail- 
able (31). At one time, round wire was prevalent, but the 
clamps necessary to support it caused the collector to jump 
and arc, so it was replaced with the figure 8 shape. 
Additional problems occurred with the figure 8 because it 
twisted and kinked when being reeled and unreeled dur- 
ing installation, and it frequently pulled out of hangers on 
curves. Consequently, the grooved type was developed and, 
together with the figure 9, has almost completely replaced 
the round and figure 8 shapes. Figure 9 and deep-grooved 
shapes are almost mandatory with a 350-MCM size and 
above, because these sizes require large splices and fit- 
tings and the widths are too large for proper tracking of 
the collector. 

The upper section of the wire, to which the support 
clamp attaches, has the same width dimension whether 
the wire is grooved, figure 8 or figure 9. Table 8.19 provides 
the necessary specifications for correct wire size selection 
(31). The most common wire is 350 or 400 MCM (both often 
called 6/0) figure 9. 

Trolley Feeder 

In order to reduce voltage drop and supply the neces- 
sary current, a feeder cable, which is uninsulated and 
stranded, is hung alongside the trolley wire. Both alumi- 
num and copper feeders are used, and their size depends 
on the load drawn by the track vehicles and the voltage 
regulation desired. Common sizes are 1,000 MCM copper 
or 1,590 MCM aluminum. Tables 8.20 and 8.21 specify 
copper feeder data (31 ). As noted in the tables, feeder can 
be purchased with a weatherproof jacket. 



Supports, Lubrication, and Turnouts 

As shown in figure 8.25, the feeder cable and trolley 
wire can be hung side by side to gain additional support 
clearance. The feeder can also be used as a messenger to 
increase the support-bracket spacing, as shown in figure 
8.26. In this configuration, a cushioning effect is provided 
for the trolley wire since the wire is free to flex under 
pressure. 

Typical brackets for supporting trolley and feeder are 
shown in figure 8.25 and 8.26. The amount of deflection or 
sag between supports can be calculated by 



D = 



3WL 2 
2T 



(8.9) 



where D = sag, in, 

W = weight, lb/ft, 
L = distance between supports, ft, 
and T = tension, lb. 

Since the figure 9 350-MCM conductor has a breaking 
strength of 12,000 lb, it can safely be tensioned to 1,200 lb, 
which is 10% of the breaking strength. This will reduce 
sag and keep the wire straight and level. The dead-end 
hooks and turnbuckles shown in figure 8.27 are used to 
install tension in the wire. 

The maximum spacing recommended for roof- 
mounted support for a semicatenary installation (fig. 8.26) 
is 20 ft. Direct suspension (fig. 8.25) spacing should be less 
than 15 ft. Table 8.22 gives support spacings on curves 
(31). When selecting proper support types and spacings, 



212 



ROUND 



0.325" 0.365" 

5 



I/O AWG 



2/0 AWG 




3/0 AWG 




4/0 AWG 



0.548" 




300 MCM 



GROOVED 




FIGURE 8 



0.106 




0.312 
1/0 AWG 



FIGURE 9 
DEEP- 
SECTION 
GROOVED 



m 



0.429 
3/0 AWG 



0.196 



0.108'i 





0.222 




O.BO 1 "*^ 

0.400" 
3/0 AWG 



•« -*- 



0.496 
350 MCM 

COPPER 



0.150' L ^ 




Figure 8.24.— Trolley-wire cross sections. 



Table 8.19.— Trolley-wire specifications 





Nominal 


Cross-sectional 


Weight 


dc resistance 
(volts drop per amp 


Minimum 


Minimum 


Elongation 


Type of wire 


size, 
AWG or 












at 20 


>C) 


tensile 
strength, 


breaking 
load, 


within 


















10 in, 




MCM 


Nominal 


Actual 


Ib/Mft 


lb/mi 


fiorV 


QorV 


psi 


lb 


% 






MCM 


MCM 


in 2 


/Mft 


/mi 






Round, hard-drawn copper, 
























97.16% conductivity 


1/0 


105.6 


105.6 


0.0829 


319.5 


1,687 


0.1011 


0.5339 


54,500 


4,518 


2.40 




2/0 


133.1 


133.1 


.1045 


402.8 


2,127 


.08021 


.4235 


52,800 


5,519 


2.80 




3/0 


167.8 


167.8 


.1318 


507.8 


2,681 


.06362 


.3359 


51,000 


6,720 


3.25 




4/0 


211.6 


211.6 


.1662 


640.5 


3,382 


.05045 


.2664 


49,000 


8,143 


3.75 




300 


300.0 


300.0 


.2356 


908.0 


4,794 


.03558 


.1879 


46,400 


10,930 


4.50 


Grooved, hard-drawn copper, 
























97.16% conductivity 


2/0 


133.1 


137.9 


.1083 


417.6 


2,205 


.07741 


.4087 


50,200 


5,437 


2.80 




3/0 


167.8 


167.3 


.1314 


506.4 


2,674 


.06380 


.3369 


48,500 


6,373 


3.25 




4/0 


211.6 


212.0 


.1665 


641.9 


3,389 


.05035 


.2659 


46,600 


7,759 


3.75 




300 


300.0 


299.8 


.2355 


907.6 


4,792 


.03560 


.1880 


44,200 


10,410 


4.50 




350 


350.0 


351.2 


.2758 


1063.0 


5,612 


.03040 


.1605 


42,800 


11,800 


4.50 


Figure 8, hard-drawn copper, 
























97.16% conductivity 


1/0 


105.6 


105.6 


.0329 


319.5 


1,687 


.1011 


.5340 


51,800 


4,294 


2.40 




2/0 


133.1 


133.1 


.1045 


402.8 


2,127 


.08021 


.4235 


50,200 


5,246 


2.80 




3/0 


167.8 


167.8 


.1318 


508.0 


2,682 


.06361 


.3359 


48,500 


6,392 


3.25 




4/0 


211.6 


211.6 


.1662 


640.5 


3,382 


.05044 


.2663 


46,600 


7,745 


3.75 




350 


350.0 


350.1 


.2750 


1 ,060.0 


5,597 


.03049 


.1610 


42,800 


11,770 


4.50 


Figure 9 deep- section, 
























hard-drawn copper, 97.16% 
























conductivity 


350 


350.0 


348.9 


.2740 


1,056.0 


5,576 


.03060 


.1616 


42,800 


11,730 


4.50 




400 


400.0 


397.2 


.3120 


1 ,202.0 


6,347 


.02687 


.1419 


41,300 


12,890 


4.50 



213 



Table 8.20.— Characteristic data for solid copper feeder cable 





Conductor 
size, 
AWG 




Section area 




Overall diameter, in 


Weight, 


Ib/Mtt 


Bare wire br 
strength, 


Baking 
lb 




cmil 


in 2 


m 2 


Bare 


Weatherproof 


Bare 


Insulated 


Hard drawn 


Annealed 


0000 




211,600 


0.1662 


107 


0.4600 


0.6163 


641 


767 


8,143 


5,320 


oon 




167,800 


.1318 


85.0 


.4096 


.5659 


508 


629 


6,722 


4,220 


00 




133,100 


.1045 


67.4 


.3648 


.5211 


403 


502 


5,519 


3,340 







105,500 


.08289 


53.5 


.3249 


.4812 


320 


407 


4,517 


2,650 


1 




83,690 


.06573 


42.4 


.2893 


.4456 


253 


316 


3,688 


2,100 


? 




66,370 


.05213 


33.6 


.2576 


.3826 


210 


260 


3,003 


1,670 


3 




52,640 


.04134 


26.7 


.2294 


.3544 


159 


199 


2,439 


1,325 


4 




41,740 


.03278 


21.2 


.2043 


.3293 


126 


164 


1,970 


1,050 


5 




33,100 


.02600 


16.8 


.1819 


.3069 


100 


135 


1,591 


880 


6 




26,250 


.02062 


13.3 


.1620 


.2870 


79 


112 


1,280 


700 


7 




20,870 


.01635 


10.6 


.1443 


.2693 


63 


NA 


1,030 


550 


a 




16,510 


.01297 


8.37 


.1285 


.2535 


50 


75 


826 


440 


NA 


Not available. 





















Table 8.21.— Characteristic data for stranded copper feeder cable 



„ .. . Number of Overall diameter, ... . . . ,. ,,,,. 

Conductor Cross-sectional area wjres jn Weight, Ib/Mft 

size in 

cmil in 2 m 2 strand Bare Weatherproof Bare Insulated 

MCM: 

2,000 2,000,000 1.571 1,014 91 1.630 1.880 6,175 7,008 

1,750 1,750,000 1.374 887 91 1.526 1.776 5,403 6,193 

1,500 1,500,000 1.178 760 61 1.411 1.661 4,631 5,380 

1,250 1,250,000 .9817 633 61 1.288 1.538 3,859 4,508 

1,000 1,000,000 .7854 507 61 1.152 1.402 3,088 3,674 

900 900,000 .7069 456 61 1.094 1.313 2,779 3,332 

800 800,000 .6283 405 61 1.031 1.250 2,470 2,992 

750 750,000 .5890 380 61 .998 1.217 2,316 2,822 

700 700,000 .5498 355 61 .964 1.183 2,161 2,650 

600 600,000 .4712 304 37 .893 1.112 1,853 2,235 

500 500,000 .3927 253 37 .813 1.001 1,544 1,894 

450 450,000 .3534 228 37 .772 .960 1,389 1,724 

400 400,000 .3142 203 19 .726 .914 1,235 1,553 

350 350,000 .2749 177 19 .679 .867 1,081 1,345 

300 300,000 .2356 152 19 .629 .817 926 1,174 

250 250,000 .1963 127 19 .574 .762 772 985 

AWG: 

0000 211,600 .1662 107 '7,19 .528 .684 653 800 

000 167,800 .1318 85.0 1 7,19 .470 .626 518 653 

00 133,100 .1045 67.4 7 .414 .570 411 522 

105,500 .08289 53.5 7 .368 .524 326 424 

1 83,690 .06573 42.4 7 .328 .484 258 328 

2 66,370 .05213 33.6 7 .292 .417 205 270 

3 52,640 .04134 26.7 7 .260 .385 163 206 

4 41,740 .03278 21.2 7 .232 .357 129 170 

5 33,100 .02600 16.8 7 .206 .331 102 140 

6 26,250 .02062 13.3 7 .184 .309 81 115 

7 NA NA NA NA NA NA NA NA 

8 16,510 .01297 8.37 7 .146 .271 51 78 

NA Not available. 

1 Sizes AWG 0000 and 000 cable are usually made of 7 strands when bare and 19 strands when insulated. 



Bare wire breaking 


Resistance, n/Mft 


strength, lb 


at 20°C, 






standard 






Hard drawn 


Soft annealed 


annealed 


87,790 


43,830 


0.005289 


77,930 


38,350 


.006045 


65,840 


32,870 


.007052 


55,670 


27,390 


.008463 


45,030 


21,910 


.010578 


40,520 


19,270 


.011753 


36,020 


17,530 


.013223 


34,090 


16,430 


.014104 


31,820 


15,340 


.015112 


27,020 


13,150 


.017631 


22,510 


10,960 


.021157 


20,450 


9,860 


.023508 


17,560 


8,765 


.026447 


15,590 


7,669 


.030225 


13,510 


6,574 


.035262 


1 1 ,260 


5,478 


.042315 


9,617 


4,637 


.04999 


7,366 


3,677 


.06304 


5,926 


2,916 


.07949 


4,752 


2,312 


.10024 


3,804 


1,834 


.1264 


3,045 


1,525 


.1594 


2,433 


1,209 


.2009 


1,938 


959 


.2535 


1,542 


761 


.3195 


1,228 


603 


.4029 


NA 


NA 


.5080 


777 


379 


.6406 



214 



>^^^^ 



TT~TT 



i— © 



5<Vie 




iHl nv\ 






KEY 



A Mine hanger 

B Dirigo spool insulator 

C Dual suspension clamp 

Z? Double yoke 

E Bulldog clamp 



F Bulldog feeder sling 

C Triple yoke 

H Mine hanger 

J Triple horizontal insulator 



Figure 8.25.— Typical trolley-wire and feeder-cable supports. 




END VIEW 



STRAIGHT LINE SPAN 



<S 55 Af^ f 




KEY 

/I Expansion bolt D Bulldog feeder sling 

B Combination clamp without boss E Combination clamp with boss 

C Mine hanger 



Figure 8.26.— Trolly-wire semicatenary suspension. 



215 




A 
B 
C 
D 

E 
F 
G 
H 
J 
K 
L 



Expansion bolt 
Mine hanger 
Dirigo spool insulator 
Dead-end hook 
Insulated turnbuckle 
Dead-end cam grip 
Feeder-wire strain clamp 
Dead-end clevis 
Feeder-wire strain clamp 
Dead-end eye 
Feeder strain clamps 



Figure 8.27.— Trolley system accessories. 



Table 8.22.— Trolley-wire support spacings on curves 



Radius of 
curve, ft 

350 and over... 

300 

250 

200 



Maximum spacing, 

ft, with deflection 

angle of 5° 

1 30 
26 
22 
17 



Radius of 
curve, ft 

150 

120 

100 

80 



Maximum spacing, 

ft, with deflection 

angle of 5° 

13 
10 

9 

7 



1 On straight lines, spacing can be increased to 30 ft where wire is 5 ft or 
more above rail. If wire is less than 5 ft above rail, the limit on inside 
construction is 20 ft. 



the weight of the trolley guards must be included. Under- 
ground coal mine regulations require guards to be posi- 
tioned wherever personnel normally work or pass under 
the uninsulated trolley and feeder wires (38). When the 
roof is uneven or too high for the trolley wire, pole 
extension brackets can be used from either the roof or rib. 

For trolley haulage outside the mine, catenary or 
direct support can be used. Simple catenary support, 
suspending the trolley wire from a messenger, works best 
for long haulage distances since 100-ft spans are possible 
on straight track. Compound catenaries, using two mes- 
senger wires and subspan catenaries are also employed. 
Semicatenary or direct suspension can be used on the 
surface, employing the same components as shown in 
figures 8.25 and 8.26 but mounting the hangers on wooden 
or metal poles or structures, and the spacing may be 
increased to 25 or 30 ft. The deflection formula (equation 
8.9) may be used for more precise calculations. 

A properly installed trolley wire looks level and 
straight, without bends, kinks, or sags. The rubbing 
surface of the metal should look polished and smooth, not 
scraped or burned bright. A graphite-based lubricant 
should be used periodically to form a smooth contact 
surface and maintain the smooth polished-brown appear- 
ance. An unlubricated, uneven trolley wire wears out 
quickly, is unsafe because of arcing, reduces power effi- 
ciency, and also wears out the collector. 



At track switches or turnouts, trolley frogs must be 
installed at the proper location to assure that the collector 
will pass on to the correct wire. Normally, a standard 10° 
trolley frog is used for any degree of track turnout. The 
frog must be positioned far enough beyond the track 
turnout that sufficient side force exists to guide the 
collector on to the correct wire. Too much side force will 
cause the collector to be pulled off the wire. The location 
for frog angles is found by determining the position where 
the collector pole exerts adequate side force on the collec- 
tor. This point occurs where the pole angle is equal to the 
track-frog angle plus 10°. 

Rails and Bonds 

Another important consideration of the trolley system 
is the resistance of the current return path: the rails and 
bonds. The specific resistance of high-carbon steel rails is 
118 fl/cmil-ft at 20° C, about 12 times that of copper. In 
table 8.23, this value is used to provide resistance values 
for mine track per rail. 

Table 8.23.— Resistance of steel rail at 20°C 



Rail weight, 
lb/yd 

40 

50 

60 


Resistance, 
WMft 

0.0250 
.0200 
.0166 
.0143 


Ri 

80 . 

90 

100 


il weight, 
Ib/Mft 


Resistance, 
WMft 

.0124 
.0111 
.0100 


70 





A major concern with rail resistance occurs at switch 
contacts and bolted rail joints. These points can have very 
large resistances, sometimes approaching an open circuit. 
Rail bonds are used to ensure a low-resistance contact or 
joint. With inadequate bonding, the return current from 
trolley system loads can stray into the mine floor or earth, 
and the stray current can cause electrical problems 



216 



throughout the mine, including nuisance tripping of pro- 
tective circuitry. 

Rail-bond cable is soft, stranded copper in sizes from 
2/0 to 500 MCM. It is attached with stud terminals or 
welds across every rail joint. Welded terminals are pre- 
ferred for permanent application on main-line haulage; 
however, the amount of heat used must be controlled in 
order to prevent steel rail recrystallization. Thermite 
welds have also been used successfully in this operation. 
Cross bonds are applied at least every 200 ft along the 
track, so that if one rail or a bond along a rail breaks, the 
current return path can be completed through the other 
rail. This practice also halves the resistance in the current 
return circuit by paralleling the two rails. Cross bonds are 
also recommended at all rail turnouts in conjunction with 
the switch points. 

All bonds are susceptible to damage by the wheels of 
derailed mine cars and hence should be located next to ties 
and secured to the tie side for protection. If possible, joint 
bonds should be placed under the rail-connecting plates. 

Bond size is usually determined by voltage drop 
rather than by current-carrying capacity. It is a general 
rule that bare conductors will carry 1 A per 5,000 cmil of 
area without excessive heating. Five times this amount 
can be carried for brief periods, and short bonds on heavy 
rail can carry 150% of the normal load current. Bond-cable 
resistances are shown in table 8.24 along with the usual 
sizes according to rail weights. The terminal resistance is 
negligible at 1 /iQ, and cross bonds are normally equal in 
length on the track gauge, typically 42 in. Joint bonds are 
usually 10 in long. The added resistance of joints, ex- 
pressed in feet of rail, can be found from the chart in figure 
8.28 {22). 

Table 8.24.— Data for rail-bond cable 



ADDED RESISTANCE PER JOINT (R), ft of rail 
20 18 16 14 12 10 8 6 4 2 





Rail-bond 


Resistance of 1 ft 


Used with rail 




size 


at 20°C, 10" 


6 fl 


size, lb/yd 


AWG: 










2/0.... 




79.35 




20- 40 


4/0.... 




49.97 




20- 80 


MCM: 










300... 




35.31 




80-100 


500... 




21.16 




80-100 



OVERHEAD LINES 

The most common method used for electric power 
transmission and distribution is overhead conductors. Al- 
though their size and detailed construction can vary 
widely, overhead powerlines normally consist of bare me- 
tallic conductors supported by insulators from some ele- 
vated structure. The conductors use air space for insula- 
tion over most of their length, while their elevation 
protects them from contact with ground objects. 

Overhead-line installations use numerous types of 
conductor arrangements and support structures in various 
combinations. Utility systems range from single wooden 
poles, carrying conductors at low voltages, to self- 
supporting steel towers bearing major transmission lines. 
Wooden polelines with or without crossarms, for example, 
may be part of a single-phase or three-phase distribution 
system with voltages of 2.3 to 35 kV. By contrast, steel 
towers often carry lines transmitting large amounts of 
power at 115 kV and up, connecting major load centers of 
a utility company grid (12, 40). Utility-owned lines are 

























30 
40 










R 

A 




R 
t 


















W ■*— 


c 


— Jw 








45 






t 




5 50 












V 










~- 55 

5 






















_i 60 
< 






















* 70 
° 80 

x 90 
S2 100 
u 














/ , 




s 




































































M& 
















A& 



















I AWG 



o 

a 



o 

CD 



H 
U 

< 



00 

000 

0000 

250 MCM 

300 
350 
400 
500 



5 10 15 20 25 30 35 40 45 50 
ACTUAL LENGTH OF BOND (L), in 

Figure 8.28.— Theoretical resistance of bonded joint. The 
proper moves to make in using this chart have been indicated 
by the small diagram in the upper central portion. Starting with 
the length (L) of the bond, move vertically to the bond capacity 
(C), then horizontally either right or left to the rail weight (W), 
then vertically to the equivalent feet of rail. 



commonly classified by function, which is related to volt- 
age. There are no utility-wide standards for voltage clas- 
sification, but the system that is typically used differs 
from the classification used in the mining industry. 

Overhead conductors are arranged in various config- 
urations to reduce line-to-line contacts due to wind, ice 
loading, or sudden loss of ice load, and may include 
different combinations of power, neutral, and static con- 
ductors. Aluminum conductors with steel reinforcement 
(ACSR) are commonly used because of their strength and 
relatively low price, but special applications may call for 
other materials such as copper (12, 40). 

The types of overhead-line installations used for mi- 
ing applications are similar to those in utility distribution 
systems. Typical are polelines to supply equipment in 
surface mining and lines feeding surface facilities related 
to mining. These lines are normally installed on single 
wooden poles and may carry only two conductors, as in 
single-phase supplies, or have up to six conductors, includ- 
ing three power, one neutral, one ground-check (pilot), and 
one static. The polelines may be relatively permanent 
installations such as those feeding plants, shops, and 
other surface facilities, and long-term pit baselines or ring 
mains. Chapter 1 includes a discussion of the poleline 
application in strip and open pit mining operations. Some- 
times, temporary poles are mounted in portable bases 
(such as concrete-filled tires) for ease of relocation, and 
these are commonly used in open pit mining operations to 
carry power into the pit. Conductors are again usually 
ACSR, but hard-drawn copper is used where blast damage 
is a problem (31). 

If these lines are not installed properly, failures from 
conductor breakage, arcing between phases, and structure 



217 



collapse can occur. Obviously, serious safety hazards and 
costly power outages can result; therefore, proper design 
and installation are important. In-depth treatment of 
overhead-line design is provided in such texts as Fink and 
Carroll (12) and Westinghouse (40), which give excellent 
detailed summaries of design, installation, and repair 
practices in overhead distribution. 

This section is intended to be a brief introduction to 
overhead-line design combined with some details of the 
wooden pole structures that are the main type found in 
mines. Overhead lines, unfortunately, have been a major 
cause of electrocutions in the mining industry; thus, an 
extensive discussion of injury prevention is also included. 

Overhead-Line Design 

The design of surface overhead lines relies as much on 
a knowledge of structure and mechanics as it does on 
electricity. The design is concerned with obtaining the 
correct size and placement of the structures that support 
the power and grounding conductors and keep them from 
damage. Obviously, the vertical weight of several single 
conductors or one multiconductor cable must be sup- 
ported. Additional vertical loading can be caused by ice 
accumulation. The height of the structures must be ade- 
quate to provide the required ground clearance consider- 
ing the amount of line sag. At the same time, the struc- 
tures must be planted firmly enough to counteract the 
force placed on them by the conductors on a steep slope. 

Tension is applied in order to install the conductors 
with the correct amount of sag, and this results in stretch. 
The stretch causes creep or elongation in the conductors 
over a period of time, which must be accommodated in the 
design. Aluminum conductors are particularly susceptible 
to creep. Another factor that must be considered is the 
effect of weather. Temperature changes cause expansion 
and contraction, which affect the amount of deflection. 
Wind causes the conductors to vibrate vertically and 
imparts a horizontal force to the structure. Calculation of 
horizontal forces is particularly important at angle points 



in the line. Additional large horizontal forces exist when a 
conductor breaks, and this factor too must be incorporated 
into the design. The vertical spacing between conductors 
must be large enough so that arcing does not occur during 
high winds or when a large accumulation of ice falls off a 
line. 

Extremely tedious calculations for catenary spans can 
result from attempting to take all of these factors into 
account. Fink and Carroll explain graphic solutions and 
describe Thomas Charts that assist in the computation 
process (12). The National Electrical Safety Code (NESC) 
(2) gives design information for ice loadings, temperature 
variations, and wind velocities for different areas in the 
United States. 

As mentioned earlier, several different types of struc- 
tures are used to support the conductors. Selection of a 
specific type depends on the terrain, accessibility require- 
ments, right-of-way availability, distribution voltage, span 
length, number of circuits, conductor size, weather, life of 
installation, availability of material, and economics. The 
types commonly used are self-supporting and guyed-steel 
or aluminum towers, steel or aluminum poles, concrete 
poles, wooden H-frames, and wooden poles. Steel and 
aluminum structures are used for high-voltage distribu- 
tion where long service life and long spans are necessary 
but are only used in some mine power systems within 
substations. Wooden poles are the most prevalent 
overhead-line support in the mining industry, so a few 
details of their design will be presented. 

Wooden poles are usually constructed from fully 
treated pine or butt-treated cedar. They are classified by 
their circumferential dimension measured at a point 6 ft 
from the butt. Consequently, the nominal ultimate 
strength is the same for all lengths and species of the same 
class. Wooden poles, classed 1 through 7 with this system, 
have the capability of withstanding the ultimate loads 
shown in figure 8.29 (12). The correct setting depth for 
various lengths is also noted in the figure. The setting 
depth is important to prevent butt "kickout," since the 
pole is primarily a cantilever column. 



GUYED POLE 

f = Safe bending stress, psi 

t = Taper, in/ft 

Moment of load = P, h,' + P 2 h 2 + P 3 h 3 

Safe moment on pole = Mof(d') 

d' = d, + tH' 



Wind on pole (W) : 



Variety 



Western cedar 
Pine 

Northern cedar 
Chestnut 



Ultimate 

fiber 

stress 

bending 



5,600 
7,400 
3,600 
6,000 



I3h(d, + d 2 ) 



Taper, 

inches circum 
per ft length 



0.38 
.25 
.63 
.40 



Average 
top diam 



d 2 -0.38h/3.!4 
d 2 - .25h/3.14 
d 2 - .63h/3.14 
d 2 - .40h/3.14 




\]/ \L w\l/\b 



UNGUYED POLE 

P = Safe load 2 in from top 
P, P 2 P 3 = Wind on wires, lb/ft 
M L = Moment of load 
M P =Safe moment on pole 
M L = P 1 h l + P 2 h 2 + P 3 h 3 
M P =PH-Wj 

r-p = '/2 max sum of adjacent spans 



Class 


Ultimate load 
2ft from top, lb 


1 
2 


4,500 
3,700 


3 

4 
5 
6 

7 


3,000 
2,400 
1,900 
1,500 
1,200 



Pole length, ft 


30 


35 


40 


45 


50 


55 


60 


65 


Setting depth, ft 


5.5 


6.0 


6.0 


6.5 


7.0 


7.0 


7.5 


8.0 



Figure 8.29.— Pole strength calculations. 



218 



Guy wires are used where necessary to assist in 
supporting the horizontal loads. Usually, 3/8-in galvanized 
steel wire tied to a log-anchored rod is used, as shown in 
figure 8.30 {12). The ratio of the guy load (L) to the 
conductor tension (T) is the same as the ratio of guy length 
(B) to the distance away from the pole base (A). 

Both wooden and steel crossarms are used on wooden 
poles. Steel crossarms give better protection from light- 
ning strokes, but they are more expensive than equivalent- 
strength wooden arms. Treated yellow pine or untreated 
Douglas fir are commonly used for crossarms. Their length 
ranges from 10 to 25 ft depending on the required conduc- 
tor spacing. In size they range from 5 by 6 in to 6 by 10 in. 
Two planks, 3 by 8 in, can be mounted on either side of the 
poles to form a crossarm for heavy conductors. 

Typical conductor arrangements and spacing for pin 
insulators are shown in figure 8.31. The dimension B is 
often determined by the span length and calculated ac- 
cording to the method given by Fink and Carroll (12). 
Because the sleet-jump failure experience, the dimension 
E should be at least 1 ft. Swing-type insulators require 
additional spacing to give a clearance for the swing. The 
clearance between the conductor and any grounded struc- 
ture must be at least 0.75 times the dry-flashover distance 
of the insulator or the "tight-string" distance under an 
8-lb wind at 60° F. v 

In addition to the conductors positioned at the insu- 
lators, overhead ground conductors or static wires are 
strung above the conductors to give lightning protection. 
Lightning protection is discussed in detail in chapter 11. 




WKfflSfflX- 



DESIGN DATA FOR GUYS 



Guy 


Ultimate 
strength, lb 


Rod 


Net area, 
in* 


Ultimate 
strength, lb 


3 /e in s.m. 


6,950 


5 /8in 


0.202 


1 1 ,000 


3 /8 in h.s. 


10,800 


3 /4 in 


.302 


16,500 


7 /i6 in s.m. 


9,350 


1 in 


.551 


30,500 


7 /i6in h.s. 


14,500 









mmm 




8-in-diameter log, 5ft long 
Weight of cone = 12,000 lb 
Allowable bearing along guy 
rod is 3 ton /ft 2 



Figure 8.30.— Guy and log-anchor calculations. 



Overhead-Line Electrocutions 



Overhead lines, whether utility transmission and 
distribution lines or part of the minesite electrical system, 
present a serious electrical shock hazard to mining per- 
sonnel. Overhead lines in and near mining operations can 
be exposed to many types of mobile equipment and even 
handheld tools. Metallic frames of such equipment, upon 
contact with energized overhead lines, can become ele- 
vated above earth potential, and simultaneous contact of 
the hot frame and ground by an individual can create a 
path through the body for lethal levels of line-to-ground 
fault current. Personnel are therefore exposed to a shock 
hazard through indirect contact with overhead powerlines. 
Although this mode of electrocution seems (at least out- 
wardly) straightforward, it has been very difficult to find 
effective means of prevention. 

Examination of mining industry statistics since 1970 
reveals that one-third of surface coal mine electrical fatal- 
ities and approximately one-third of all electrical fatalities 
in metal-nonmetal operations are directly attributable to 
the indirect contact of overhead lines (25). The majority of 
these accidents involved mobile equipment, and the haz- 
ard can exist anywhere that high-reaching equipment 
operates near overhead lines (1). 

Trucks commonly involved in overhead-line contacts 
are highway-legal end-dump tandems, triaxles, and trac- 
tor trailers. They can contact overhead lines, and their 
frames become subsequently energized, through their 
beds being raised or driven into lines. Victims normally 
bridge lethal potentials when stepping from the cab onto 
the ground or by operating external controls. Mobile 
cranes, which present a substantial line-contact hazard in 
other industries, find various uses around a mine site. 
They range from large, solid-boom, high construction 



C' ^ 



X 



jSl 



U- 



^T 



■b-I 



Pole top 



r»- 



B--1EN- 

S. It 



-fl 



^w- ! d 



&- 



wU 



"W' 



Two arm 






Single arm 



Voltage 


Spacing 


A 


B 


D 


F 


23 kV 


3'-0" 


5'-0" 


2'-6" 


3'-0" 


34.5 kV 


3'-4" 


5'-0" 


3'-0" 


3-4" 


46 kV 


4'-0" 


7'-0" 


5'-0" 


5-0" 


69 kV 


6'-0" 


8'-0" 


6'-0" 





Figure 8.31.— Typical arrangements and pin-insulator spac- 
ings on wooden poles. 



cranes to smaller hydraulically powered units with re- 
tractable booms. Lines can be contacted by the boom or 
hoisting cable, and in both cases workers around the crane 
have the greatest shock-hazard exposure. Mobile drilling 
rigs are susceptible to overhead-line contact because of 
their masts, which can be raised or driven into the lines. 
Operators are the most likely victims, bridging potential 
gradients while operating drill controls. 



219 



In response to the problem, a detailed investigation 
has been made into these mining accidents, into preven- 
tion methods used by utility companies and other indus- 
tries, and into various additional methods that might 
reduce electrocutions from indirect contact with overhead 
lines (27). From this effort, typical hazardous mining 
locations with overhead lines were identified and several 
recommendations were established to reduce the associ- 
ated hazards. 

The listing of areas and situations that pose the 
greatest overhead-line hazard is important since it shows 
the target areas for application of recommended solutions. 
These locations can be divided into two groups: mining 
surface facilities and active excavations in surface mines. 

Loading and dumping facilities, including stockpiles, 
loading bins or hoppers, material transfer points, and 
adjacent areas, yards and roads, are hazardous overhead- 
line locations for truck operation. Some factors contribut- 
ing to the risk are operator unfamiliarity with the dump, 
use of a temporary dump point, and fluctuations in the 
edge or height of a stockpile. Trucks and cranes can easily 
be exposed to line hazards near various mine plant areas 
such as mineral processing, storage, handling installa- 
tions, refuse dumps, and settling ponds. Construction sites 
may or may not be near permanent mining facilities but 
often present hazards involving construction cranes and 
preexisting overhead lines. 

Overhead lines traversing active surface mine work- 
ings present potentially dangerous situations. The fatali- 
ties that have occurred in these areas were from lines 
other than pit power distribution. Hazards exist primarily 
over mine benches as well as access and haulage roads. 
Although not responsible for electrocutions in the past, pit 
power distribution can create a hazard when overhead 
lines are used, such as for strip-mine base lines. 

The recommendations to reduce these hazardous sit- 
uations include isolating overhead lines from mobile 
equipment, modification of overhead lines, use of protec- 
tive devices, and safe work practices, each of which will be 
discussed in the following paragraphs. 

Overhead-Line Isolation 

It is the responsibility of the power engineer in a 
surface mine to assess the overhead distribution system 
with regard to the movement of mobile equipment and to 
ensure that wherever possible overhead lines are isolated 
from travel routes. This may seem an obvious course of 
action, but previous accidents have shown that correctable 
hazardous situations are often allowed to exist at mining 
operations. 

Where there is frequent dump-bed truck traffic, lines 
must be restricted from dump sites and approach or exit 
roads. A safety margin of at least 100 ft should be allowed 
outside normal truck routes. This would allow for limited 
truck movement beyond the route to account for mechan- 
ical problems, bed cleaning, backups and temporary dump 
sites. Roads leading away from dump locations should not 
be crossed by lines for at least 250 ft beyond the dump site, 
since beds may not be completely down as trucks leave the 
area. This distance would give additional time for the bed 
to lower or for the driver to recognize the condition. 

Construction cranes that remain stationary while 
operating at a project site can be positioned so that line 
contact cannot occur at any position. Cranes that travel 
during operation will require barriers around hazardous 
areas. When a safe distance from overhead lines is being 



determined, contact by hoist cables and swinging loads 
should be considered. 

One situation is which line isolation may not be 
feasible is where the lines supply power to a surface 
facility or a nearby installation. In order to eliminate bare 
overhead conductors in these situations, some alternate 
method must be used to supply power. One alternative for 
permanent installations is underground cable. Cables in 
conduit or directly buried are suitable for lines entering 
plants, dump facilities, shops, supply yards, and support 
buildings. Cables similar to those found in mine power 
distribution, such as MPF and SHD types, are used for 
buried applications. 

Cables present a safe alternative to bare overhead 
conductors in areas where high-reaching equipment must 
travel. Underground service removes line exposure com- 
pletely, but overhead cable with pole support may be 
preferable because of cost, ground conditions, or expected 
installation life. In either method, the cable should com- 
pletely span the hazardous area, or its purpose is defeated. 
These cable runs should continue for a short distance 
beyond the hazard area to allow for equipment extensions 
protruding beyond area limits. 

Overhead lines traversing active surface mine work- 
ings present a hazard to high-reaching equipment. 
Whether they are preexisting utility lines or part of mine 
power distribution, hazards can result for trucks and drills 
on benches or on haulage and access roads. The removal of 
these lines from the work area is the most direct solution. 
This may involve the permanent relocation of a utility line 
over a proposed open pit or a temporary rerouting of a line 
about a strip operation. Elimination of overhead lines in a 
pit power-distribution system would probably involve re- 
placement of cable. Operations such as strip mines can 
and commonly do use all-cable distribution with good 
results, provided that proper cable-handling techniques 
and equipment are used (27). Open pit operations nor- 
mally use overhead distribution to switchhouses in the pit 
and shielded trailing cables to mobile equipment. How- 
ever, none of the fatal accidents examined were due to 
contact of these overhead distribution lines. In large open 
pit mines, overhead distribution is the most practical 
because of the long distances and cable protection require- 
ments, but where frequent equipment operation poses a 
contact hazard, cable may be more desirable. 

When rerouting lines around surface mine work ar- 
eas, all aspects of the operation should be considered, 
including surface clearing, reclamation, access roads, and 
haulage roads, as well as actual mining activities. A 
safety margin should again be provided beyond normal 
work areas to account for occasional abnormal truck 
traffic, excavator booms, and similar situations. 

Contact with overhead lines can also be avoided by 
removing the equipment operation from the hazardous 
area instead of moving the lines. Although this should be 
a very effective method, sometimes equipment movement 
is necessary: for instance, access cannot be restricted for 
cranes in supply yards or trucks in dump areas. However, 
where lines traverse active surface mine workings, equip- 
ment could be kept out of any contact-hazard area. Limit- 
ing access to lines can be the only economic alternative for 
a very small strip operation, which may be unable to 
sustain the cost of relocating even a small overhead 
distribution line. Nevertheless, any efforts to restrict mo- 
bile equipment must be carefully planned and imple- 
mented so as not to hamper normal operations or antago- 
nize the work force. 



220 



It may be possible to restrict high-reaching equipment 
from some permanent surface facilities. Where this is 
possible, it provides an effective and less costly alternative 
to relocating overhead lines, so long as normal operations 
are not hindered. Restriction can be accomplished by 
posting the area, or using barriers such as steel crossbars, 
which allow only low vehicles (cars and small trucks) into 
the area. Provisions can easily be included to allow occa- 
sional entrance of higher equipment. 

One option for the operation is to avoid the hazard by 
leaving the overhead-line right-of-way undisturbed. How- 
ever, this option can result in a loss of resource as well as 
a disruption in the continuity of mining. The right-of-way 
may involve forfeiting only a single pass, as in a contour 
strip operation, but may seriously affect the mine layout if 
a large-area strip mine is traversed by a major transmis- 
sion line. In order for a contour mining operation with an 
overhead powerline across the projected path to continue 
through the right-of-way but not mine below the lines, the 
towers or poles beyond the pit width limits would have to 
be guyed. The cables could then be removed or lowered 
into trenches, and all large equipment would be trammed 
or walked over the right-of-way. The lines would then be 
replaced, and mining operations would resume on the far 
side of the overhead lines. 

Exploration drilling commonly requires operation in 
unfamiliar surroundings, often under minimal supervi- 
sion. However, drill sites may usually be relocated to avoid 
overhead-line hazards. 

Overhead-Line Modification 

Solutions discussed prior to this point isolate over- 
head lines from mobile equipment to reduce the change of 
contact. There are modifications to existing overhead lines 
that can substantially reduce hazards without resorting to 
the extreme measures stated earlier. Such techniques are 
important because many cases will arise where an opera- 
tor cannot eliminate overhead-line hazards nor limit ac- 
cess to them. 

Overhead-line heights must never be less than the 
minimum mandated by Federal regulations (38). These 
heights are extracted from the NESC for driveways, haul- 
ageways, and railroads, and 15 ft is stipulated as the 
minimum height for any high-voltage power line (2). Table 
8.25 lists the NESC standards that cover most overhead 
lines in mining, while table 8.26 provides the required 
minimum distances for higher voltages. 

Some hazards can be reduced by raising some over- 
head lines above the NESC minimums. Where dump-bed 
truck traffic is a concern, lines over roadways could be 



raised to clear most dump-bed units without extensive 
support structure. A line height of 45 ft would place lines 
above most highway-legal dump-bed trucks, even with 
their beds fully raised, and would also clear most drills 
and cranes when they are in transit with their booms and 
masts lowered. If necessary, it is possible to raise lines to 
more than 65 ft using single wooden-pole supports. How- 
ever, the line heights attainable depend upon line spans, 
cable sag, and surrounding terrain, but in most cases 45 ft 
is an achievable height. 

Another line modification that lends itself to road 
crossings is the guarding of power conductors by effec- 
tively grounded conductors. If it can be ensured that any 
accidental contact with power conductors will be simulta- 
neous contact with grounded conductors, a line-to-ground 
current will probably be provided. This reduces current 
flow through an equipment-ground contact and increases 
the chance of rapid fault clearing by circuit protective 
devices. However, several grounded conductors will be 
necessary to ensure simultaneous contact and may make 
this method impractical because of cost. Under these 
circumstances, rubber guarding may be used on overhead 
lines at hazardous crossings. 

Utilities will often supply electricity to a mining 
facility substation by running a branch overhead line from 
their lines. If the branch line creates a contact hazard on 
or around the mining property, a disconnect switch should 
be provided external to the utility system and upstream 
from any contact-hazard area. Should the need arise to 
work in close proximity to these lines, power could be cut 
with no disturbance to other utility customers. Discon- 
nects that are quickly accessible from mine work areas 
would also encourage deenergization prior to work about 
lines, but this depends upon ownership of the lines, 
availability of qualified personnel to cut power, and utility 
policy. 

Protective Devices 

Devices exist that attempt to reduce overhead-line 
hazards either by insulation from line potentials or warn- 
ing of overhead-line proximity. Representative of the insu- 
lation method are insulated boom cages and insulating 
load hook links; proximity warning devices are intended to 
indicate the presence of energized conductors. Most de- 
vices are directed primarily toward protection of mobile 
cranes but do have other applications. 

An insulated boom cage is an enclosure or guard 
mounted on and electrically isolated from the boom or 
mast to be protected. If the boom is moved into an 
energized overhead line, the insulated cage makes initial 



Table 8.25.— Minimum vertical conductor clearances as specified by the NESC, applicable to mining and mining-related 

operations 



Criteria 



Nature of surface 

underneath wires, 

conductors, or cables 



Open supply line 
conductors, ft 



750 V to 
15 kV 



15 to 
50 kV 



Locations where wires, conductors or cables cross over. . 



Locations where wires, conductors, or cables run along and 
within the limits of highways or other road right-of-ways but do 
not overhang the roadway. 



Track rails of railroads (except electrified railroads using 28 30 

overhead trolley conductors). 

Roads, streets, alleys, parking lots subject to truck traffic 20 22 

Other land traversed by vehicles, such as cultivated, grazing, 20 22 

forest, orchard, etc. 

Roads in rural districts 18 20 



221 



Table 8.26.— Minimum distances from overhead lines for 
equipment booms and masts (38) 



Nominal 
powerline 
voltage, kV 

69 to 114 

115 to 229 

230 to 344 



Minimum 

distance, 

ft 

12 
15 
20 



Nominal 
powerline 
voltage, kV 



345 to 499.. 
500 and up . 



Minimum 

distance, 

ft 

25 
35 



contact and prevents the boom from becoming energized. 
The device only protects covered areas and cannot easily 
guard hoisting ropes, and its effectiveness also depends on 
the integrity and surface condition of the insulators used 

(i). 

During crane hoisting operations, workers steadying 
or directing a load from the ground are in an extremely 
hazardous situation should an overhead line be contacted, 
as they are commonly in contact with both the ground and 
load. Insulating links can be used to isolate loads from the 
crane hoisting rope and are placed between the load hook 
and the hoist rope. The links are constructed of a dielectric 
such as resin-impregnated fiberglass. 

A proximity warning device indicates by a visual or 
audible alarm the proximity of equipment extensions to 
energized overhead powerlines (15). Unlike cages and 
links, these devices attempt to prevent equipment-line 
contact, and the operation is theoretically independent of 
human judgment, at least so far as indication of powerline 
presence is concerned. Ideally, such devices alert an oper- 
ator if the protected equipment extension enters a prede- 
termined zone about a power conductor. Several types of 
proximity warning devices are available in the United 
States, all operating on the principle of electrostatic-field 
detection (15). The electrostatic field about a group of 
overhead conductors is primarily a function of their volt- 
age and geometry. The units generally operate by moni- 
toring 60-Hz electrostatic fields, amplifying, rectifying, 
and measuring the signal, and then activating an alarm at 
some preset signal level. The sensor used may be short and 
effectively a point sensor, which will create a spherical 
detection area, or a distributed sensor spanning the length 
of a protected extension. The type, number, and location of 
these sensors greatly affect the operation of a proximity 
warning device. Proximity warning devices operate as 
intended under many circumstances, but their reliability 
can be compromised by a complex array of factors. These 
limitations can be grouped into those arising from opera- 
tional principles of electrostatic-field detection and those 
which are due to the design of individual devices. 

The concept of a device to alert equipment operators to 
possible overhead-line contacts has great merit, but given 
the inconsistent operation of currently available devices, 
they should only be applied with full recognition of their 
limitations. Dangerous conditions can exist where work- 
ers place too much faith in a warning device or ignore it 
due to previous unreliable operation. Proximity warning 
devices are best applied only as a supplement to other 
overhead-line contact safety measures. Boom cages and 
insulating load links also have sound theories of operation 
but problems in implementation, and major drawbacks 
stem from flashover due to insulator surface conditions. 
The use of any warning or insulating technique does not 
relieve the operator from the responsibility of maintaining 
the minimum line-equipment clearances stated earlier. 



Safe Work Practices 

Any attempt to reduce overhead contact hazards at a 
mining operation must also involve the development of 
safety awareness within the work force. Training of per- 
sonnel in safe operation of mobile equipment near over- 
head lines will complement any other safety method and, 
in some cases, may be the only effort necessary for 
preventing indirect-contact electrocutions. The following 
recommendations include guidelines for work near over- 
head lines, some passive-warning techniques, and safety 
training of personnel. 

Before work is done near high-voltage overhead lines, 
the areas in question should be thoroughly examined by 
supervisory personnel and workers to determine the pres- 
ence of any overhead-line hazards. All overhead lines 
should be considered energized unless an authorized rep- 
resentative of the line owner indicates otherwise. If the 
lines are utility owned, the utility should be contacted for 
assistance with planning safe operating procedures for the 
project. Equipment should be operated only by a compe- 
tent, experienced, qualified operator, and the operations 
should be observed by a reliable worker, watching for 
maintenance of minimum clearances and unsafe condi- 
tions. This observation should be the worker's designated 
and only task. Another competent worker should be des- 
ignated to direct the equipment operator, and only this 
worker should give directions. Standard signals should be 
agreed upon and used. Booms, masts, beds, and so forth 
should be in a lowered position when equipment is in 
transit, and minimum legal clearances should be main- 
tained. If minimum clearance cannot be provided, the 
overhead lines in question should be deenergized and 
visibly grounded. 

The following procedures should be followed if an 
energized overhead line is contacted. If contact was mo- 
mentary and no lines are down, a calm and experienced 
crew member should be certain that the equipment is no 
longer in contact and should then assign members of the 
crew to check for injuries among the work party, to 
administer first aid if necessary, such as basic life support 
and cardiopulmonary resuscitation, and to send for an 
ambulance immediately, to notify supervisory personnel, 
to check for dangerous equipment damage, and to secure 
the area for possible accident investigation. If contact is 
made and maintained, a calm and experienced crew mem- 
ber should instruct personnel aboard the equipment to 
remain in place and not to contact the ground, then have 
the operator move equipment out of contact if possible. 
Crew members should be assigned to keep all other 
personnel clear of the area, including equipment, hoisted 
loads, and fallen lines, to notify appropriate mine super- 
visory personnel or utility to have lines deenergized, and 
to send for an ambulance if needed. The crew should not 
contact any victims still in contact with energized equip- 
ment. When victims can be rescued safely, the crew should 
administer first aid, move equipment to a safe position, 
check for damage, and secure the area for possible acci- 
dent investigation. 

Investigation of past fatalities shows clearly how 
essential it is for workers to be familiar with these 
procedures, and the importance of regular training in 
cardiopulmonary resuscitation (CPR). 

Passive-warning techniques, including signs, stickers, 
posters, and line indicators, should be highly visible and 
in color to draw worker attention. They should be to the 



222 



point and simple to understand. Signs in hazardous areas 
should be large enough to be easily read from approaching 
equipment and should warn operators well in advance of 
the danger. 

Paragraphs 48.25 through 48.28, and 48.31, 30 CFR, 
mandate the initial training and periodic retraining of 
mine personnel with respect to the occupational hazards of 
mining. High-voltage overhead-line safety should be in- 
cluded in this training. New employees at surface opera- 
tions are often laborers assisting on or about mobile 
equipment, and in their initial training they must be 
alerted to the danger presented by overhead lines. Hazards 
specific to the mining facility in question should be 
brought out in initial training and retraining, as well as in 
the hazard training required for workers assigned to new 
jobs, particularly new equipment operators. Frequent re- 
views of safe practices regarding overhead lines would be 
advisable for all operators of high-reaching equipment, 
regardless of the minimum legally required training. 
Particularly important is the review of safety guidelines 
with crews about to begin operations with exposure to 
overhead lines. Familiarizing supervisory personnel with 
safety guidelines and company policies is also essential if 
they are to be competent in directing the work force under 
hazardous conditions. 



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224 






CHAPTER 9.— PROTECTIVE EQUIPMENT AND RELAYING 



Even the best designed electrical systems occasionally 
experience faults and overloads, or disturbances that 
cause abnormally high currents. These currents can exist 
in the ground system or in the phase conductors. Wherever 
the occurrence, the situation is likely to precipitate a 
hazard to either equipment or personnel. 

Of the basic design criteria that underlie all mine 
power systems, three are of critical importance in protec- 
tive equipment and relaying: adequate interrupting capac- 
ity, current-limiting capacity, and selective system opera- 
tion. The first two provide protection to the system during 
a disturbance, while the third is designed to locate the 
problem, then minimize its effect. In chapter 7, current 
limiting and selective relaying were designated as two 
prime purposes of grounding. It was shown that ground- 
fault currents can be limited by inserting a resistance in 
series with the neutral conductor. However, not much has 
been presented about selective system operation, other 
than its need. Protective circuitry and protective relaying 
are the tools behind selective system operation and are the 
main topics of this chapter. 

The protective circuitry associated with the power 
system consists of transducers, relays, and switching ap- 
paratus. Its role of safeguarding personnel and equipment 
can be effected manually or automatically. An instance of 
manual utilization would be removing power from a sys- 
tem portion for maintenance. An example of automatic 
operation would be a situation in which protective cir- 
cuitry first senses then clears each hazardous current 
resulting from a disturbance. As might be expected, the 
process of clearing is disconnecting the affected circuit 
from the power source safely and as quickly as possible, 
with minimum interference to the system balance. In 
other words, protective circuitry must isolate a malfunc- 
tion at a given location with minimum damage to circuits 
and equipment and minimum operation downtime. The 
function of protective circuitry to provide detection and 
isolation is termed selective relaying. 

All the devices that comprise the protective circuitry 
in the mine power system thus play a vital role in safety. 
In fact, protective circuitry is probably the most important 
component of the power system and forms a major portion 
of all power equipment. For example, a switchhouse, 
which has the principal function of protection, is simply a 
complex of protective devices. 

The basic concepts of overloads and faults are intro- 
duced in chapter 4. Although the removal of destructive 
overloads is important, the main concern is the clearing of 
faults, since their occurrence can be catastrophic. 

Because of the preponderance of cables, cable shield- 
ing, and grounded equipment in mine power systems, 
line-to-neutral faults are the most common, and most of 
these are arcing with relatively short length and con- 
trolled distance. Ground-fault current is predominantly 
limited by neutral grounding resistors, whereas in other 
industrial applications, ground-fault currents are often 
limited by fault impedance. 

Line-to-line and three-phase faults can also occur, as 
when a mobile machine severs a cable during a runover. 
Extremely large line currents can result, which can be 
limited in the mine system only by transformer and 
line-conductor impedances. System components, such as 
couplers, cables, transformers, bus bars, and disconnect 
switches, must be capable of withstanding the momentary 



mechanical and thermal stresses created by the flow of 
fault current through them. Interrupting devices, such as 
circuit breakers, must be able not only to withstand these 
momentary fault-through stresses, but to interrupt or 
terminate these anomalous currents. 

The maximum magnitude of possible fault currents 
existing in line conductors must be known in order to 
select adequate ratings of protective equipment. Indeed, 
this knowledge is required to coordinate protective- 
circuitry operation for the entire complex. It may also be 
necessary to know the minimum sustained fault current 
that is available in the system in order to determine the 
sensitivity requirements of the current-responsive protec- 
tive devices. These fault magnitudes, both maximum and 
minimum, are usually estimated by calculation, and the 
equipment is selected using the calculated results. 

Because of the many hazards that can occur, the system 
must be capable of detecting overloads, short circuits Qine 
faults), undervoltage, and ground faults, as well as any 
compromise in grounding-conductor continuity. With the use 
of resistance grounding in mine power systems, the protec- 
tive relaying or sensing device associated with ground faults 
or zero-sequence currents is usually handled separately from 
that for line faults causing only anomalous positive-sequence 
or negative-sequence currents. In addition, the relaying for 
overloads may be separate from that for faults. Except for 
fuse applications, the sensing devices for each function will 
normally cause the activation, or tripping, of the same 
circuit-interrupting device no matter what the protection 
requirements are for an individual location. The sensing 
devices may be an integral part of the interrupting appara- 
tus or be separated from it and connected only through 
control wiring. 

This chapter builds upon the material covered in 
chapter 4, beginning with the main protection compo- 
nents, switching apparatus and sensing devices. Basic 
relay connections, relay terminology, and different kinds 
of protection follow. Finally, typical assemblies and com- 
binations of protective circuitry are discussed. Essentially, 
this chapter sets the stage for chapter 10, where fault 
calculations, device sizing, and coordination are outlined. 

SWITCHING APPARATUS 

A switching apparatus is defined as a device for 
making (closing), breaking (opening), or changing connec- 
tions (6). 1 There are three basic types of apparatus in this 
classification: switches, circuit breakers, and fuses. 

All switching devices are given certain design ratings, 
which are a measure of the electrical stresses they can 
withstand (6). Obviously, the ratings must be correlated 
with the intended use or duty. A listing and definition of 
these ratings follows but is restricted to those terms 
having direct application to the development of the topic in 
this and the subsequent chapter. Further concepts will be 
added in the discussion of transients and overvoltages in 
chapter 11. 

1. Voltage. The maximum nominal system voltage at a 
specified frequency (usually line-to-line for ac devices) on 
which the device may be installed. 



1 Italicized numbers in parentheses refer to items in the list of references 
at the end of this chapter. 



225 



2. Continuous current. The maximum continuous cur- 
rent that the apparatus may carry. 

3. Short-circuit current. Usually, the maximum cur- 
rent the device is capable of interrupting. This may be 
further qualified by an interrupting-current or 
interrupting-capacity rating. 

4. Close-and-latch or momentary current. The maxi- 
mum short-circuit current that the device can withstand 
during the few cycles after the fault occurs without expe- 
riencing severe mechanical damage. 

The ratings of switching apparatus are based on the 
maximum possible values of fault currents. To help visu- 
alize the importance of these ratings, consider that a 
three-phase fault has occurred on a power system. Figure 
9.1 illustrates the resulting line current versus time, 
created by the flow of energy from the source or sources to 
the fault (7). This asymmetrical waveform is made up of 
two components: dc and a symmetrical ac. At any instant 
after the fault occurs, the total fault current equals the 
sum of the two. The dc component decays to zero in a short 
time, with the total current gradually changing from 
asymmetrical to symmetrical. 

Switching-apparatus ratings, as a measure of the 
stresses involved during faulting, are based on the sym- 
metrical rms value. Asymmetry is accounted for by taking 
the basic symmetrical value and applying multiplying 
factors. These concepts are presented in detail in chapter 
10. However, figure 9.1 does provide a useful visualization 
of rating magnitudes, and these will be discussed in 
following sections, along with each switching device. 



Total asymmetrical current 



dc component 



Symmetrical ac 




TIME 



Figure 9.1. —Typical system fault current. 



Wh 



-Contacts closed, circuit energized 



Load or short circuit 



ARCS AND CIRCUIT INTERRUPTION 

After a switching apparatus receives a message that 
circuit current is to be interrupted, the device proceeds 
through definite steps to terminate the current (4). These 
are illustrated in figure 9.2. Under normal operation, the 
contacts of the apparatus are closed, current flows through 
the interface, and the outgoing circuit is thus energized. 
To terminate current, the contacts begin to separate and 
an electric arc is drawn. 

The arc is composed of free electron and free positive- 
ion flow, as shown in figure 9.3. To initiate this arc, free 
electrons and/or free positive ions must exist between the 
contacts. Their availability depends upon the following 
environmental conditions: 

• In air or gas, the conductive elements are generated 
prior to the initiation of the arc by radiation and cosmic rays, 
which knock off electrons from neutral gas molecules. 

• In a liquid such as oil, the conductive elements exist 
as impurities. 

• In a vacuum, they can be emitted from the cathode 
by a high-strength electric field with the process known as 
high field emission. 

The last case can add free electrons to any environment. 
Even though the voltage between the cathode and anode is 
low immediately after separation, the free electrons are 
attracted to the anode, and the positive ions toward the 
cathode. The electron flow accounts for about 90% of arc 
current (4). 

If the voltage across the arc remains large enough, the 
movement of charge between the contacts initiates the 
mechanisms that can increase and sustain the arc. This 



/ Contacts parting, arc drawn 
1-711-71 between contacts 



Load or short circuit 



•Contacts open, arc extinguished 



-0 LZh 



Load or short circuit 
deenergized 



Figure 9.2.— Steps in circuit interruption. 



Region B 
positive column 



Region A 




Electrons emitted in great number 
thermionically or from cathode spot 

Figure 9.3.— Arc between two contacts. 



226 



again depends upon the environment. In a gas, the free 
electrons can collide with neutral gas molecules, producing 
additional free electrons and positive ions, termed ionization 
by collision. In any atmosphere, the collision of heavy posi- 
tive ions on the cathode produces heat, which can augment 
field emission in low-melting-point materials such as copper, 
creating intense electron discharge from a small area, called 
a cathode spot, and can cause thermionic emission in high- 
melting-point substances, such as carbon, where electrons 
are boiled out by high temperature. 

Once the arc is established, processes must be brought 
into play to extinguish it. In general, the greater the arc 
current and the higher the voltage of the circuit, the more 
difficult is the problem of arc extinction. The situation is 
easier in ac systems than in dc systems because the 
current waveform passes through zero in ac systems. 
However, the arc can restrike when the voltage rises again 
if the ionic conditions across the contacts permit. For dc, 
the arc is readily maintained because a normal cur- 
rentzero does not exist. 

Whatever the extinction process, the switching device 
can open the circuit successfully, provided that the current 
to be interrupted is within the rated value. However, if the 
switching device is required to terminate a current well 
above the design value, the arc between the parting 
contact may not extinguish or may continue to restrike, 
and the apparatus could be destroyed by the gas pressure 
built up within it (5). 

When a device is designed to interrupt fault current, 
it is often called an interrupting device; otherwise, it is 
commonly called a switch and is designed only to open and 
close a circuit. While some switching apparatus are in- 
tended to serve only one of these functions, others can do 
both. 



SWITCHES 

A switch has exactly the same definition as switching 
apparatus, with the qualification that it is a manual 
device (6); in other words, its operation is a normal or 
intended occurrence on the power system. The switch 
types common in mine systems are the disconnect and the 
load interrupter. Both have the prime function of isolating 
outgoing circuits from the power source. 

A disconnect switch is not intended to interrupt 
circuit current and can be operated only after the circuit 
power has been removed. Interlocks must be provided to 
prevent manual operation under load, and latches may be 
needed to prevent opening from the stresses resulting from 
fault circuits. Consequently, disconnect switches do not 
have an interrupting rating; but beyond a continuous- 
current rating, they may need a momentary-duty or close- 
and-latch rating for handling fault-through currents. 

An interrupter or load-break switch differs from a 
disconnect in that it has an interrupting rating. The 
device has the capability of terminating currents that do 
not exceed the continuous-current rating, although this is 
not its normal operation. Interrupter switches usually 
have a quick-make, quick-break mechanism, which pro- 
vides a fast switch-operation speed independent of the 
handle speed. The illustration in figure 9.4 shows a 
three-pole device; the mechanism on the right side of the 
connecting shaft provides the fast operation. Some units 
can be motor driven, thus allowing remote or automatic 
operation. In most mining applications, load-break 
switches need a close-and-latch rating. Where interlocks 




Figure 9.4.— Load-break switch. (Courtesy Line Power Manufac- 
turing Corp.) 



are not employed, this rating indicates the margin of 
safety when the switch is closed into a faulted circuit. 

Switches are normally used as disconnects in mining 
systems regardless of their ratings; in fact, some States 
require load-break switches with interlocks for all discon- 
necting applications. These interlocks cause interruption 
of source power prior to contact separation, and the 
operation is usually performed through the ground- 
monitoring circuitry. Load-break switches, used in con- 
junction with fuses, are employed as interrupters in cer- 
tain circumstances. 



CIRCUIT BREAKERS 

A circuit breaker is primarily an interrupting device, 
but in some cases it is also used as a switch (6). A circuit 
breaker can be defined as a device designed to open and 
close a circuit manually and to open the circuit automat- 
ically at a specific current level without injury to itself 
when properly applied. It is available as a single pole, 
double pole, or triple pole. Manual operation, be it me- 
chanically or electrically actuated, is again intended 
where the circuit current is not in excess of rated contin- 
uous current. Automatic operation is dictated by a system 
abnormality, such as a fault or an overload. In this case, 
the device may be called upon to interrupt current in 
excess of the rated continuous current. 

Circuit breakers in the role of interrupting devices must 
be used with sensing devices to perform their intended 
function. In medium-voltage and low-voltage mining appli- 
cations, the operation may be internally controlled by self- 
contained current-responsive elements, external protective 
relays, or a combination of both. In high-voltage situations, 
the sensing devices are always separate, with interconnec- 
tions only through control wiring. 

Circuit breakers can generally be broken into two 
classifications: those intended for systems over 1,000 V, 
and those for 1,000 V and below. Devices in the first class 
are called power circuit breakers, while the second class is 
divided into power circuit breakers and molded-case cir- 
cuit breakers. Following mining standards, circuit break- 
ers for systems below 661 V are called low voltage; for 661 
to 1,000 V, medium voltage; and above 1,000 V, high 
voltage. It should be noted that IEEE Standards define 
above 1,000 V to 72,500 V as medium voltage and below 
1,000 V as low voltage. Low- voltage and medium-voltage 
circuit breakers are usually considered together and can 






227 



find ac and dc service. High-voltage breakers involve only 
ac circuits. The next paragraphs look at typical apparatus 
and operation. 

CIRCUIT BREAKERS FOR LOW AND MEDIUM 
VOLTAGE 

The term air circuit breaker is often used when 
referring to molded-case and power circuit breakers de- 
signed for low-voltage and medium-voltage systems (7). 
Air circuit breakers employ the simplest method of inter- 
rupting current: extinguishing the arc in normal atmo- 
sphere by increasing its length (4). Several different pro- 
cesses can be used to force the arc to lengthen. 

lb illustrate one arc-lengthening technique, consider 
figure 9.5, where two circuit breaker contacts, a and b, 
have just separated. The horn-like arrangement of the 
contacts shown in the figure can be considered an arc 
chute, which is a barrier that confines, cools, and extin- 
guishes the arc (5). By the ionization of the air between the 
contacts, an arc is drawn and heat is generated. The arc 
extinction action then commences; this is also called 
deionization because it serves to reduce the free electrons 
and positive ions in the gas (4). Air currents, created by the 
heat and confined by the arc chute, force the arc upward to 
form a loop. Electromagnetic forces within the loop further 
encourage the lengthening. As a result of cooling by 
radiation or convection, the longer arc requires a higher 
arc voltage to sustain current flow, and thus, the arc is 
extinguished. 

As noted earlier, arc interruption in an ac circuit 
occurs much more easily than in a dc circuit. All voltages 
and currents in an ac system go through cyclic changes, 
and consequently, ion-producing effects for the arc are 
variable too: falling as current becomes smaller, ceasing at 
current zero (4). Deionizing effects in the arc chute remain 
steady. To take advantage of this situation, circuit break- 
ers for ac systems are often designed around the minimum 
voltage required to establish a cathode spot. Because there 
is no natural current zero in dc systems, the circuit 
breaker must force the current to zero. For this to happen, 
the arc voltage must be greater than the system voltage 
(14). An enormous amount of heat can be generated in all 
circuit breakers while the arc exists, and an important 
function of the circuit breaker assembly is to dissipate this 
heat safely. 

The foregoing simple arc-lengthening technique 
works well for 240-Vac applications. Conventional practice 
is to use a single-pole breaker for 120 Vac and a double- 
pole breaker for 240-Vac single-phase circuits. The latter 
employs one pole of the circuit breaker in series with each 
power conductor. 

For circuits 250 V and above, the direct arc-lengthen- 
ing approach is not enough; special arc chutes, quenchers, 
or deionizing chambers are needed to assist in arc termi- 
nation (5). Figure 9.6 illustrates one approach, where the 
arc is forced into metallic barriers by magnetic attraction 
and broken into a series of smaller arcs. Each of these arcs 
is subjected to lengthening, cooling, and the problem of 
reestablishing a cathode spot if low-melting-point materi- 
als are used (4). Another approach is depicted in figure 9.7. 
Because the arc establishes its own electromagnetic field, 
an external magnetic field can enhance arc lengthening. 
The process is termed magnetic blowout, and breakers 
using this principle are called air magnetic. Coils carrying 
the circuit current in series with the arc can provide the 



Arc rises on horns 




V b 
Arc drawn here 
Figure 9.5.— Extinguishing arc by increasing the length. 



a b 




Metal barriers 
Arc 

Arcing contacts 



Main contacts 



Figure 9.6.— Metal-barrier arc chute assists in arc deioniza- 
tion. 



Barriers of 
insulating material 




Arcing contacts 

Figure 9.7.— Insulated-barrier arc chute used with magnetic 
field. 



228 



magnetic field. As shown in the figure, the magnetic field 
forces the arc into insulated barriers or fins, creating 
further lengthening; recombination and cooling at the 
barrier surfaces accelerates deionization (4). 

In dc mine power circuits below 660 V, air-magnetic 
breakers are used extensively, especially on trolley sys- 
tems. With very few exceptions, molded-case breakers are 
employed for ac circuits below 1,000 V. In addition, 
molded-case units are often used to protect low-voltage dc 
face equipment. 

Molded-Case Circuit Breakers 

The molded-case circuit breaker is the most explicit 
example of interrupting apparatus with self-contained 
current-responsive elements. It is defined as a breaker 
that is assembled as an integral unit in a supporting and 
enclosing housing of insulating material (5). Depending 
upon the amount of protection desired, these devices can 
sense internally and then clear undervoltage, overcurrent, 
and short-circuit conditions. Some tripping elements, that 
is, the actual components that cause the contacts to start 
separating, are also externally accessible through control 
wiring. Hence, other circuit protection can be added. 
Except for some power circuit breakers of low-voltage and 
medium-voltage design, all the circuit breakers that will 
be discussed in this chapter rely solely on outside infor- 
mation to perform their prime function. Molded-case ap- 
paratus will be presented first so that many important 
terms can be introduced. 

The application of molded-case circuit breakers in 
mining began in the 1950's with the conversion from 
low-voltage dc power distribution to ac power distribution 
and face rectification, expanding further with the trend 
toward ac face equipment. In fact, Wood and Smith (21) 
have attributed the introduction of low-height, solid-state 
rectifier units in underground mines (which permitted the 
use of ac distribution) to molded-case circuit breakers, 
citing the lack of high-speed dc circuit breakers of the 
proper height as the previous limitation. Molded-case 
breakers placed between the transformer and the rectify- 
ing bridge lowered the height limitation to that of the 
transformer, allowing a unit design complementary with 
the mining environment. 

The largest mining application is trailing-cable pro- 
tection in underground face areas. The breakers are lo- 
cated in power centers and provide cable protection on 
each outgoing circuit, as required by 30 CFR 75.900, in 
addition to functioning as switching devices. The typical 
molded-case breakers, however, are not designed for repet- 
itive switching. Mining use subjects them to many more 
operations than found in other industries, and regular or 
standard circuit breakers generally cannot hold up to the 
stress. Several manufacturers, recognizing this problem, 
have produced a special line of mine-duty molded-case 
breakers, which have stronger construction to withstand 
the punishment of mine use. 

Except for external adjustments, molded-case devices 
sometimes do not allow field maintenance; many are 
sealed to prevent tampering. Although some manufactur- 
ers offer a complete line of replacement components, 
repairs other than an exchange of easily removable parts, 
such as arc chutes or trip units, should be made only by 
qualified repair facilities. This is critical, given the impor- 
tance of the molded-case circuit breaker in personnel 
protection. 



All component parts of these circuit breakers are built 
into one insulated housing, the molded case. These parts 
are the operating mechanism, arc extinguishers (arc 
chutes), contacts, trip elements, and terminal connectors, 
as shown in figure 9.8 (19). Additional accessories may be 
included. 

The molded case is made of a glass polyester or 
similar synthetic material that combines ruggedness and 
high dielectric strength with a compact design. Each type 
and size of molded case is assigned a frame size or 
designation for easy identification. This coding, loosely 
based on an old Underwriters' Laboratories standard, 
refers to a number of breaker characteristics, including 
maximum allowable system voltage, maximum allowable 
continuous current, interrupting capacity, and the physi- 
cal dimensions of the molded case. Several trip units may 
be available for a particular frame size, so a specific 
assembled breaker may have a lower continuous-current 
rating than the current designation of the frame. Table 9.1 
lists the continuous ratings considered to be standard for 
mining service. The currents in parentheses are the lower 
current settings available in that frame size from certain 
manufacturers. Unfortunately, manufacturers have vary- 
ing design criteria and hence size their units to dissimilar 



Table 9.1. —Ratings for mining-service molded-case circuit 
breakers 

Frame size, ' A Continuous-current ratings, 2 A 

100 100 (70, 50, 30) 

225 225 (175, 150, 125, 100) 

400 400 (300, 225, 175, 150, 125, 100) 

600 600 (500, 400) 

800 800 (600) 

1,200 1,200 (1,000, 800, 600) 

1 Regular-duty breakers also available in 1,600-, 2,000-, and 2,500-A 
frames. 

2 Currents in parentheses are lower settings available in the frame size. 



Molded case (frame) 

\ Operating mechanism 




Arc extinguishers 

Contacts 



Trip elements 



Figure 9.8.— Molded-case circuit breaker components. 
(Courtesy Westinghouse Electric Corp.) 



229 



specifications. For example, a 225-A, 600-V breaker sup- 
plied from two separate manufacturers may have different 
physical dimensions so that direct interchanging is diffi- 
cult, if not impossible. 

The circuit breakers rated in table 9.1 are generally 
available as two-pole or three-pole units at 600 Vac or 300 
Vdc, but only as three-pole devices at 1,000 Vac. The 
two-pole breakers are intended for dc face equipment or 
single phase ac applications. By convention, one pole is 
used for each ungrounded conductor in a circuit (5). 

The arc chutes define the interrupting-current capac- 
ity of the assembly in conjunction with the insulating and 
heat-dissipation properties of the molded case. The chutes 
assist arc deionization by the principle discussed for figure 
9.6. They are also termed arc extinguishers or arc quench- 
ers by some manufacturers. The breaker case must be 
mounted vertically with the arc chutes at the top for 
correct arc-extinction operation. 

Circuit breakers designed for 1,000 V and below are 
capable of clearing a fault faster than those constructed 
for high voltage (6). The contacts often begin to part 
during the first cycle of fault current, and consequently, 
the breaker must be capable of interrupting the maximum 
allowable first-cycle asymmetrical current. Thus, for lower 
voltage breakers, the close-and-latch and interrupting 
ratings are usually the same, a characteristic not found 
with high-voltage breakers. The rating of these units is 
carried out on a symmetrical basis, so multipliers account- 
ing for the dc offset need not be applied as long as the 
system X/R ratio does not exceed 6.6 (6) (see chapter 10). 
Table 9.2 lists typical interrupting ratings versus the 
system voltages for mine-duty circuit breakers; the ac 
system values are based on the symmetrical rating. Some 
manufacturers offer both standard-duty and high- 
interrupting-capacity breakers for mining service. The 
table values presented parenthetically indicate the supe- 
rior construction, which incorporates sturdier contacts 
and mechanism plus a special high-impact molded casing. 

Table 9.2 shows that typical molded-case circuit 
breakers constructed for 1,000- Vac mine systems have only 
a 10,000-A symmetrical interrupting rating. This presents 
a concern, as available short-circuit currents on high- 
power 1,000-Vac systems can be greater. Instances include 
longwall mining equipment, which needs a power-center 
capacity of 1,500 kVA or more. To overcome the problem, a 
manufacturer has introduced molded-case breakers with a 
24,000-A asymmetrical interrupting rating at 1,000 Vac 
and continuous-current ratings of 600, 800, 1,000, or 1,200 
A. The asymmetrical rating is used to provide more 
flexibility for designing the breaker into power systems. 

The function of the operating mechanism of a typical 
molded-case circuit breaker is to provide a means of 
opening and closing. It is a toggle mechanism of the 
quick-make, quick-break type, meaning that the contacts 
snap open or closed independent of the speed of handle 



movement. The breaker is also trip-free; that is, it cannot 
be prevented from tripping by holding the breaker handle 
in the ON position during a fault condition. In addition to 
indicating whether the breaker is ON or OFF, the 
operating-mechanism handle indicates when the breaker 
is tripped by moving midway between these positions. To 
reactivate the tripped breaker, the handle must first be 
moved from the central position to OFF, which resets the 
mechanism, and then to ON. This distinct trip point is 
particularly advantageous where molded-case breakers 
are grouped, as in a power center, because it clearly 
indicates any faulty circuits. 

The function of the trip elements is to trip the 
operating mechanism in the event of prolonged overload or 
short-circuit current. Two common types of trip elements 
are used in mining, magnetic and thermal magnetic. 
When the circuit being protected involves portable or 
trailing cables, the thermal-magnetic combination is 
strongly recommended and is mandated by some States. 

The magnetic trip protects against short circuits, and 
an electromagnet in series with the load current provides 
the trip action (19). This type of short circuit is actually a 
line-to-line or three-phase fault on ac, or a line-to-line fault 
on dc systems. When a short occurs, the high fault current 
causes the electromagnet in the breaker to attract the 
armature, initiating an unlatching action, which in turn 
causes the circuit to open (fig. 9.9). The action takes place 
within 1/2 s (usually within 1 cycle or 16 ms), instanta- 
neously tripping the breaker. Since tripping takes place 
with no intentional delay, the magnetic trip is often called 
the instantaneous-trip element. Screwdriver slots, located 
on the front of the trip unit, are used in adjusting the 
sensitivity (fig. 9.10A). By law, the maximum setting is 
established by the protection of the minimum conductor 
size in the circuit (16-17). Table 9.3 lists these maximum 
settings applied to trailing cables. Figure 9.10B illustrates 
a family of time-current curves resulting from the adjust- 
able range; to the left or below each curve, the breaker will 
not be tripped magnetically. Typical instantaneous-trip 
ranges versus frame sizes for mining-service breakers are 
given in table 9.4. Note that this is not a rigorous listing, 
since some manufacturers will provide any desired trip 
range with most frame sizes upon request. 

The other common molded-case breaker type is the 
thermal-magnetic variety. In addition to providing short- 
circuit protection, the thermal-magnetic breaker also 
guards against long-term current overloads existing 
longer than roughly 10 s, by incorporating thermal trip 
elements (fig. 9.11). The thermal action is accomplished 
through use of a bimetal strip heated by load current (19). 
The strip consists of two pieces of metal bonded together, 
each with a different coefficient of thermal expansion. A 
sustained overload causes excessive heating of the strip, 
resulting in deflection of the bimetal, which in turn causes 
the operating mechanism to trip the breaker. Because the 



Table 9.2.— Interrupting-current ratings 1 versus system voltage, amperes 



Frame size, A 240 Vac 

100 18,000 (65,000) 

225 25,000 (65,000) 

400 42,000 (65,000) 

600 42,000 (65,000) 

800 42,000 (65,000) 

1,200 42,000 (65,000) 

1 Parenthetical ratings are for typical premium-duty circuit breakers. 

2 Actual dc interrupting current dependent upon system inductance. 



480 Vac 



600 Vac 



1 ,000 Vac 



300 Vdc 2 



14,000 (25,000) 


14,000(18,000) 


10,000 


10,000 (20,000) 


22,000 (35,000) 


22,000 (25,000) 


10,000 


10,000(20,000) 


30,000 (35,000) 


22,000 (25,000) 


10,000 


10,000(20,000) 


30,000 (35,000) 


22,000 (25,000) 


10,000 


10,000(20,000) 


30,000 (35,000) 


30,000 (25,000) 


10,000 


10,000(20,000) 


30,000 (35,000) 


30,000 (25,000) 


10,000 


10,000 (20,000) 



230 



Table 9.3.— Maximum instantaneous-trip settings 







Maximum 






Maximum 


Conductor 


allowable 


Conductor 


allowable 




size 


instantaneous 
setting, A 




size 


instantaneous 
setting, A 


AWG: 






AWG: 






14. 




50 


1 .... 




1,000 


12. 




75 


1/0. 




1,250 


10. 




150 


270. 




1 ,500 


8... 




200 


3/0. 




2,000 


6... 




300 


4/0. 




2,500 


4... 




500 


MCM: 






3... 




600 


250 to 


2,500 


2... 




800 









Magnetic element 



Load 




Magnetic element closes gap and 
opens contacts on short circuit 




Latch 



^Contacts closed 
Latched 



^ Contacts open 



Tripped 

Figure 9.9.— Magnetic-trip relay. 



Table 9.4.— Commonly available magnetic-trip ranges for 
mining-service molded-case breakers 



Frame size, 
A 

100 

225 

400 

600 

800 

1,200 



Magnetic-trip 
range, A 



Range of allowable 
conductor sizes 



50- 180 

150- 500 

300- 700 

500-1,000 

300-1 ,000 

500-1 ,000 

800-1 ,600 

500-1,500 

900-3,000 

750-1 ,500 

1,000-2,000 

1 ,500-3,000 

2,000-4,000 

1,500-3,000 

2,000-4,000 

2,500-5,000 



14-10 AWG 

10-4 AWG 

6-3 AWG 

4-1 AWG 

6-1 AWG 

4-1 AWG 

2-2/0 AWG 

4-2/0 AWG 

1 AWG - 500 MCM 

2-2/0 AWG 

1-3/0 AWG 

2/0 AWG - 500 MCM 

3/0 AWG - 500 MCM 

2/0 AWG - 500 MCM 

3/0 AWG - 500 MCM 

4/0 AWG - 500 MCM 




Low 

Intermediate 

High 



ABC 
CURRENT 



A Adjustment knob B Characteristics 

Figure 9.10.— Adjustable instantaneous setting. 



bimetal deflection is dependent upon current and time, 
the thermal unit provides a long-time delay for light 
overloads and a fast response for heavy overloads. A 
representative current-time curve for the thermal unit 
alone is shown in figure 9.12A; later in this chapter, it will 
be described as an inverse-time characteristic. In compar- 
ison, figure 9.12B shows the circuit breaker response when 
both thermal and magnetic trip elements are incorpo- 
rated. The shaded area for each curve represents a toler- 
ance between the minimum and maximum total clearing 
time. 

The thermal-magnetic unit shown in figure 9.11 is 
ambient-temperature sensitive. Assuming the circuit 
breaker, cable, and equipment being protected are in the 
same ambient temperature, the circuit breaker trips at a 
lower current as the ambient temperature rises in corre- 
spondence to safe cable and equipment loadings, which 
vary inversely with ambient temperature (19). Thermal- 
magnetic trip elements are available that automatically 
compensate for ambient-temperature variations. The am- 
bient compensation is obtained through an additional 
bimetal strip, which counteracts the overload bimetal. 
Such trip units are recommended whenever the protected 
conductors and the circuit breakers are in different ambi- 
ent temperatures (19). 

Most mining-service molded-case breakers with 225-A 
frame sizes and above have interchangeable trip units. For 
straight magnetic elements these allow different 
instantaneous-trip ranges per frame size. However, 
thermal-magnetic units can be used to establish a lower 
continuous-current limit for the breaker. The National 
Electrical Code (13) is used as a guide to define the current 



Magnetic element 
/ 



Load 




Load K 




Contacts open 
V 7 , 



Contacts closed 



Tripped 

Figure 9.11.— Thermal-magnetic action of molded-case cir- 
cuit breaker. 



1,800 



10 





1a 




















8 











135 500 

CURRENT, % 

of breaker rating 

A Thermal only 



Figure 9.12.— Time-current characteristics for thermal- 
magnetic molded-case circuit breakers. 



60 


''^A — I — 


Thermal 
action 


0.016 


i|jb 


- Magnetic 
action 








; , 






250 4,000 




CURRENT, % 
of breaker rating 


B 


Thermal magn 


etic 



231 



at which the long-time-delay thermal element must ini- 
tiate the circuit-clearing operation and specifies a point 
that is 125% of the rated equipment or conductor ampac- 
ity. As seen in figure 9.12A, the circuit breaker will take 
no action below this current. Hence, the thermal portion 
defines the continuous-current rating of the breaker, speci- 
fied as 100% at 40° C for conventional (non-compensating) 
thermal-magnetic elements. Obviously, the thermal element 
current rating cannot exceed the frame rating. Because of 
the connection, some manufacturers recommend that the 
continuous current through the breaker be limited to 80% of 
the frame size. This topic will be continued in chapter 10. 

Electromechanical magnetic and thermal-magnetic 
trip elements have been replaced by solid-state compo- 
nents in some molded-case breakers. Although the solid- 
state counterparts may become popular in the future, they 
have not yet achieved wide acceptance in the mining 
industry. Nevertheless, these breakers are discussed in 
chapter 14. 

The last basic breaker components are the terminal 
connectors. Their function is to connect the circuit breaker 
to a desired power source and load. They are usually made 
of copper and must be constructed so that each conductor 
can be tightened without removing another. The terminal 
connectors shown in figure 9.8 are for direct connection of 
one cable connector per terminal. Many molded-case 
breakers also have provisions for threaded-stud terminals. 
These studs can be used not only for connection of more 
than one conductor per terminal, but also for breaker 
mounting. It should be noted that the type of terminal 
used on a breaker may change its heat dissipation proper- 
ties and thus lower its interrupting rating. 

In addition to the basic components, several accesso- 
ries are available, of which the most common are the 
terminal shield, shunt trip, and undervoltage release 
(UVR). Terminal shields protect personnel from accidental 
contact with energized terminal connections and are sim- 
ply plates that shield (guard) the terminals. The other two 
accessories are used to trip the operating mechanism. 

A shunt trip is employed to trip a circuit breaker 
electrically from a remote location. It consists of a 
momentary-rated solenoid tripping device mounted inside 
the molded case that activates when control power is 
applied across the solenoid coil. The magnetic field created 



by the solenoid moves a plunger, which in turn activates a 
trip bar. At the same time, a series cutoff switch removes 
power to the solenoid coil, preventing it from burning up 
under continuous load. A typical shunt-trip assembly is 
shown in figure 9.13. The shunt trip can remotely trip the 
breaker but cannot remotely operate it. To reclose the 
breaker, the handle must first be moved to the reset 
position, then to the ON position. 

The purpose of the UVR is to trip the breaker when- 
ever control voltage to the UVR falls below a predeter- 
mined level, usually 35% to 70%. This device is also 
mounted inside the breaker frame and consists of a spring 
and a solenoid. The spring is cocked or precharged by the 
operating mechanism when the breaker is closed and is 
held in the cocked position by the solenoid after closure. If 
the voltage drops below the required level, the solenoid 
releases the spring, causing the circuit breaker to trip. The 
breaker cannot be turned on again until the voltage 
returns to 80% of normal. 

The importance of the shunt trip and UVR is far 
ranging, as they allow the protection capabilities of circuit 
breakers to be extended. The molded-case breaker alone 
can provide overload and short-circuit protection in an 
outgoing circuit. The UVR adds undervoltage protection; 
in fact, undervoltage protection is normally required at 
most breaker locations. Note that undervoltage protection 
is required for all equipment, but it is not required on all 
circuit breakers as long as all equipment downstream 
from the breaker has undervoltage protection. The under- 
voltage protection provided by a UVR is actually "loss- 
of- voltage" protection since the dropout level is well out- 
side the recommended operating range of most motors (see 
chapter 6). Through a specific combination of relays and 
sensing devices, additional types of protection can be 
applied through shunt or UVR tripping. With a shunt trip, 
the relay completes the circuit between the control-power 
source and the solenoid coil. When a UVR is used, the 
relay removes the control voltage across the solenoid coil. 
This circuitry will be discussed in detail later in the 
chapter. 

The molded-case circuit breaker is the most widely 
used breaker in mining, even though its employment is 
restricted to low-voltage and medium-voltage systems. The 
principal application is on ac, where it provides high 



B 











Figure 9.13.— Shunt-trip (A) and undervoltage-release (8) accessories. (Courtesy General Electric Co.) 



232 



interrupting capacity for short circuits in minimum space. 
On ac or dc systems, it is often the first protection device 
called upon to handle electrical problems existing on 
trailing cables and mining machinery. A clear understand- 
ing of the construction and rating of these breakers is 
required to assure adequate protection. The operating 
characteristics must be closely matched with those of the 
trailing cable to minimize hazards to personnel. 

Power Circuit Breakers 



listed values, frame sizes are available up to 6,000-A 
continuous ac current and 12,000-A continuous dc current 
(5). These frame sizes are rated to carry 100% of the 
continuous-current rating inside enclosures at 40°C. In 
power breakers with low current ratings, arc interruption 
can utilize arc-chute arrangements similar to those used 
in molded-case breakers. The full air-magnetic arrange- 
ments described for figure 9.7 are employed for high- 
current-interruption power breakers. 



Some mining-industry engineers have found that 
molded-case circuit breakers cannot handle the available 
short-circuit currents in certain low-voltage applications, 
such as the outgoing dc circuits of trolley rectifiers and dc 
face equipment. The low-voltage power circuit breaker 
provides an alternative in these cases. 

Power circuit breakers for applications of 1,000 V and 
below are of open construction assembly with metal 
frames. They are designed to be field maintained under 
planned periodic inspection, and all parts are accessible 
for ease of maintenance, repair, and replacement (6-7). 
The design enables higher endurance ratings and greater 
repetitive-duty capabilities than are available from 
molded-case devices. However, power circuit breakers are 
intended only for service inside enclosures with "dead- 
front" construction, that is, not accessible to unauthorized 
personnel. 

Electromechanical units are available for long-time 
tripping, but mechanical-displacement dashpot types are 
normally used for this function and provide the same 
overcurrent protection as does the bimetal thermal trip- 
ping in molded-case breakers. Although long-time charac- 
teristics are not adjustable with bimetal strips, the dash- 
pots allow the long-time-delay "pickup" current and 
operation time to be changed. This extends the capabili- 
ties of the power circuit breaker over the molded case by 
providing not only short-circuit but also overload tripping 
adjustments, thereby allowing a broader range of applica- 
tions (7). Low-voltage power circuit breakers are available 
with or without direct-acting instantaneous units and 
with or without long-time-delay units. Furthermore, most 
manufacturers offer three different separately adjustable 
long-time-delay operation bands as well as three different 
short-time-delay operation bands. As with molded-case 
breakers, power breakers are available with either shunt- 
tripping or UVR units or both. Solid-state devices are also 
manufactured for all tripping arrangements. 

Some typical ratings for low-voltage power circuit 
breakers are provided in table 9.5 (7). In addition to these 

Table 9.5.— Some typical ratings for low-voltage power circuit 
breakers 



Ac system Rated 
nominal maximum 
voltage, V voltage, V 



Frame 
size, A 



3-phase short-circuit 
current rating, 
symmetrical, A 



Range of 

trip-device 

current ratings, A 



600. 



480. 



635 



508 



225 


14,000 


40- 225 


600 


20,000 


40- 600 


1,600 


42,000 


200-1 ,600 


2,000 


42,000 


200-2,000 


3,000 


65,000 


2,000-3,000 


4,000 


85,000 


4,000 


225 


22,000 


40- 225 


600 


30,000 


100- 600 


1,600 


50,000 


400-1,600 


2,000 


50,000 


400-2,000 


3,000 


65,000 


2,000-3,000 


4,000 


85,000 


4,000 



HIGH-VOLTAGE CIRCUIT BREAKERS 

The power circuit breakers used in high-voltage min- 
ing applications include air-magnetic, oil, minimum-oil, 
and vacuum types. Vacuum circuit breakers or VCB's are 
by far the most popular because of their small size and 
high efficiency. Oil circuit breakers or OCB's once were the 
most common, but their use has dropped substantially in 
recent years, since the interrupting sizes needed for min- 
ing are not available. Air-magnetic types are normally 
limited to surface installations. The next few paragraphs 
will examine typical apparatus ratings, and then the 
operation of oil, minimum-oil, and vacuum types will be 
described; air-magnetic breakers are excluded as their 
operation is the same as that presented previously for 
lower voltage breakers. 

Typical Ratings 

The typical nominal voltage ratings corresponding to 
nominal system voltages are 4,160, 7,200, and 13,800 V, 
with 23,000 V used in some strip mines. The system 
portions of interest are obviously ac. Common continuous- 
current ratings are 400, 600, 800, 1,200, and 2,000 A. The 
majority of mine systems do not call for current greater 
than 600-A continuous, which has become the most used 
rating. 

Interrupting and close-and-latch ratings are very im- 
portant high-voltage parameters (6). For low-voltage and 
medium-voltage circuit breakers, the two ratings are usu- 
ally the same. As high- voltage circuit breakers rarely 
terminate current flow until a few cycles after the first- 
cycle peak, the close-and-latch rating must be higher than 
the interrupting rating. A typical interrupting rating for 
high-voltage circuit breakers found in mining is 12,000-A 
rms symmetrical, while the typical close-and-latch rating 
is 20,000-A rms asymmetrical. The asymmetrical close- 
and-latch rating is often found by multiplying the sym- 
metrical interrupting rating by 1.6 (see chapter 10) (6). 

High-voltage circuit breakers can also be given an 
interrupting-capacity class, which is an identifying group- 
ing rather than a rating. It is expressed in megavoltam- 
peres, such as 250, 350, 500, and 750 MVA. The interrupt- 
ing capacity is related to the interrupting-current rating 
by (5) 

MVA = V3 kV^kA^a (9.1) 

where MVA = interrupting capacity, MVA, 

kV rated = rated system voltage, kV, 
and kAra^j = rated rms interrupting current, kA. 

Oil Circuit Breakers 

Even though their popularity has been dropping, 
OCB's are still used extensively in surface installations, 



233 



especially substations. The common type of construction is 
the dead tank, shown in figure 9.14A. This steel tank is 
partly filled with oil and has a cover with porcelain or 
other composition bushings or insulators through which 
the conductors are carried (4-5). The breaker contacts are 
located below the bushings and are bridged by a conduct- 
ing crosshead supported by a lift rod. In most designs, two 
contacts and the crosshead provide two interruptions per 
pole. The majority of OCB's in mining have three such 
poles in one tank. The tank has an insulated liner to 
prevent the arc from striking the tank walls. The entire 
assembly is oiltight; a vent with oil-separating properties 
permits the escape of any gases generated but prevents the 
escape of entrained oil. 

Arc interruption in high-voltage circuit breakers em- 
ploys the cathode-spot phenomenon combined with arc 
lengthening and deionization of the arc path. In the case of 
the OCB, oil is vaporized as an arc is established between 
the parting contacts, and this produces a bubble around 
the arc. The gases within the bubble are generally not 
conducive to ionization, but in most modern OCB's, an 
oil-filled insulating chamber surrounds the parting con- 
tacts (fig. 9.145). When the moving contact is lowered, the 
gas generated by the arc portion within the chamber forces 
oil out through the chamber throat (4). The blast of oil 
comes into intimate contact with the arc, accelerates the 
cooling and ion recombination process (fig. 9.14C), and 
carries away available ions. A different arc-chamber ap- 
proach is shown in figure 9.15. Here the chamber throat is 
made of laminations so that during interruption, the oil 
can move radially into the arc path. This is sometimes 
termed a turbo action. In high-interrupting capacities, the 
gases developed within the chamber can be used to blast 
oil horizontally across the arc path. Whatever the specific 
design, the chambers are intended to contain the devel- 
oped high gas pressures and reduce any pressure on the 
main oil tank (5). After being effectively cooled, the 
generated gases are allowed to pass through the vent into 
open air. 

The result of OCB construction and operation is a very 
effective arc interrupter. However, beyond availability, 
there are inherent disadvantages that discourage use of 
OCB's (4-5). The oil presents a fire hazard, particularly if 
the tank is ruptured because of unexpected pressure; this 
has led some States to prohibit OCB application in under- 
ground coal systems above 10,000 V. The oil is bothersome 
to handle and creates maintenance problems including 
cleanliness problems. Finally, the inertia of the heavy 
operating mechanism severely limits operational speed, 
causing a time delay in opening the arc. Despite these 
problems, other advantages, which are discussed in chap- 
ter 11, still make the OCB desirable to many industry 
engineers. 

When used underground, the physical size of three- 
pole units usually limits the interrupting capacity to 100 
MVA or less, with continuous-current ratings of 400 A. 
The operating mechanism on these small OCB's is typi- 
cally spring-gravity and manual-reset; a handle-driven 
mechanism (quick break, quick make) is used to close the 
breaker manually while at the same time automatically 
tensioning an opening spring. With the breaker engaged, 
the spring becomes armed, allowing a shunt-trip or UVR 
device to trigger the breaker opening by releasing the 
spring. A motor-driven system is also available to close the 
breaker, but the tripping method is the same. The motor- 
driven OCB's can thus be electrically engaged as well as 



Moving contact 




Oil level 

Arc chamber 
Insulating lift rod 




Stationary contact 

Moving contact 

Insulating chamber 
Throat 



Figure 9.14.— Construction and operation of dead-tank OCB. 



r~\ 





• Moving contact 



Oil flows into throat 
between laminations 



Figure 9.15.— Turboaction arc chamber for OCB's. 



tripped. Larger OCB's such as those used in substations 
are typically motor driven. 

Minimum-Oil Circuit Breakers 

Minimum-oil circuit breakers, also termed low- 
volume oil or live tank, enclose each pole in its own 
small-diameter tank (5). In modern versions, the tank is 
made of insulated high-strength, high-resistance material, 
and the top and bottom covers are high-dielectric-strength 
insulators (fig. 9.16). Contacts consist of a movable vertical 
rod and a stationary contact in the tank bottom. Oil 
volume is about 1 L, and the top surface of the oil is at 
atmospheric pressure. Arc extinguishing is assisted by oil 
blast, and resulting gases are vented to outside air. The 
operating mechanism can be either manual-reset and 
spring-trip or motor-reset and spring-trip. Some typical 
ratings of these breakers are listed in table 9.6. 

The arrangement of a three-pole minimum-oil unit 
with moving contacts mechanically interconnected results 
in a smaller overall package than comparable dead-tank 
breakers. The smaller mass of moving parts (operating 



234 



Moving 
contact 



Stationa 
contact 




-Operating 
mechanism 



Figure 9.16.— Cross section of minimum-oil breaker. 
Table 9.6.— Typical minimum-oil circuit breaker ratings 



Rated voltage, 
V 


Interrupting capacity, 
MVA 


Continuous current, 
A 


5,000.... 


173 
478 
680 


1,000 


15,000 


1,000 


25,000 


630 







mechanism and rods) enables higher operating speeds, 
while the advantages of oil interruption are maintained. 
However, the low volume of oil is such that after about five 
operations, the oil level must be checked. Even though 
oil-level indicators are available, this can create a main- 
tenance problem in mining. 

Vacuum Circuit Breakers 

With all the circuit breaker types covered so far, a 
gaseous atmosphere exists between the parting contacts. 
The gas is ionized by many processes and thus provides 
free electrons, which move to the anode, and positive ions, 
which are attracted to the cathode (4). As the positive ions 
arrive at the cathode, they can cause thermionic or high- 
field emission of electrons, which has a negative effect on 
arc interruption. Almost all these phenomena cease to 
exist if the gas between the breaker contacts is removed; in 
other words, if the arc is drawn in a vacuum. For this 
reason, vacuum is considered an extremely good medium 
for switching, and circuit breakers have been developed to 
take advantage of this feature. Figure 9.17 shows a sketch 
of a VCB, again with one pole. The assembly is sometimes 
called a bottle. 

The main advantages of VCB's are 

• Interruption usually occurs at the first zero current; 

• There are no blind spots in their interrupting 
range; 

• They have extraordinarily long life; 

• They are relatively maintenance free; and 

• Recovery of dielectric strength (between the parting 
contacts) following interrupting is extremely fast. 



Moving 
contact 

\ 



Insulation 

1 



Metal 
bellows 



^M 



U 



Stationary 
contact 



High vacuum 
Figure 9.17.— Cross section of VCB. 



These all result from the fact that the vacuum totally 
discourages ionization. 

An important aspect of VCB's is in the long service. 
For instance, if a unit fails to clear a short circuit beyond 
its interrupting range, but another unit down the line 
does, the exceeded VCB can be employed again up to the 
full rating without difficulty. Because of their efficient 
ratio of size to capacity, they are extremely well suited to 
underground mining use. Their interrupting capacity for 
large currents is such that they can be utilized anywhere 
on high-voltage distribution, usually without reservation. 
This flexibility has made the VCB the most popular 
high-voltage interrupter for distribution systems in min- 
ing today. 

Added to these advantages is the fact that the VCB does 
not have any physical orientation problems. This is a consid- 
erable constraint with OCB's, where the tanks must always 
be vertical. Vertical placement of VCB bottles is sometimes 
necessary, however, to minimize dust accumulation. 

Ironically, the high efficiency of vacuum interrupters, 
which has favored their wide application, is the same 
property that can lead to severe transients. If care is not 
taken with VCB installation, switching transient-related 
problems can occur throughout the mine electrical com- 
plex. A detailed discussion of this important problem is 
deferred until chapter 11 because of related phenomena. 

The operating mechanism, which includes the mount- 
ing structure for the vacuum bottles, is an important 
factor in proper VCB operation. As a result of the small 
contact travel distance, usually on the order of 1/4 in, four 
criteria are mandatory: 

1. Rugged construction to withstand the shock and 
stress of equipment movement; 

2. A firm, smooth closure motion to prevent contact 
bounce; 

3. Forceful opening of contacts in the case of contact 
welding; and 

4. Clean, smooth opening motion to prevent contact 
bounce and subsequent arc restriking. 

In most cases, manufacturers rely on a spring-reset and 
spring-trip mechanism to meet items 2 through 4, and 
figure 9.18 illustrates one approach. The closing opera- 
tion, also termed resetting or reclosing, may be manually 
or motor driven. The trip solenoid can be a shunt-trip or 
UVR device, and in some cases, both are used. 

In VCB applications, the compact size of the operating 
mechanism and mounting structure has made possible a 
substantial reduction in overall power-equipment dimen- 
sions. Manufacturers have even incorporated a disconnect 



235 



Contact opening 
spring (extended) 



Contact 
pressure 
spring 

(compressed 

Main contacts 
(closed) 




Bounce latch (disengaged) 

v.. 

Trip latch 
engaged ) 



V 

v Trip coil 
(deenergized) 



Reset latch 
(engaged) 



Mechanism 
rec losing lever 



Main contacts closed 



Contact opening 
spring (relaxed) 



Contact 

pressure 

spring 

(relaxed) 




Bounce latch (engaged) 



^-Trip coil 
(energized) 
Motor 



Main contacts -— ^== 1 

(open) ^Cp/ ! ff / 

_. . . . v ^s£W — Mechanism 

Reset latch reclosing lever 

(disengaged) 

Contact travel just completed after tripping 



Contact opening 
spring (relaxed) 



Contact 

pressure 

spring 

(relaxed 



Main contacts 
(open 



Bounce latch (disengaged) 

Trip coil 
(deenergized) 

J 




Motor turns lever 
to rec lose contacts 



Reset latch 
(disengaged ) 



Main contacts open, ready for reclosing 

Figure 9.18.— Operating mechanism for vacuum interrupter. 

(Courtesy McGraw Edison) 



switch in their designs (fig. 9.19). The operating mecha- 
nism for the switch is mechanically interlocked with the 
circuit breaker mechanism. If the switch is opened when 
the breaker is closed, the interlock trips the circuit 
breaker prior to switch-contact parting. 



FUSES 

The fuse is the simplest and oldest device for inter- 
rupting an electrical circuit under short-circuit or 
excessive-overload current (5, 7). Fuses are installed in 
series with the protected circuit and operate by melting a 
fusible link. The response is such that the greater the 
current, the shorter the time to circuit opening, that is, an 
inverse-time characteristic. Fuses may be used in ac or dc 
circuits, and there is such variation in their time-current 




Figure 9.19.— VCB assembly incorporating a load-break 
switch. (Courtesy Ensign Electric) 



characteristics that they are suitable for many special 
purposes. While circuit breaker contacts rely on external 
sensing, the fuse acts as both the sensing device and the 
interrupting device. Unlike circuit breakers, fuses are 
"one-shot," as their fusible element is destroyed in the 
circuit-protection process. Fuses are available with 
interrupting-current ratings up to 200,000-A symmetrical 
rms, much higher than the capacity of circuit breakers. 
Fuses are also available with current-limiting abilities to 
provide maximum protection for all circuit components. 

Fuses are normally classified as low voltage or high 
voltage: The low-voltage types are intended for service in 
systems 600 V and below, while the high-voltage varieties 
are suitable for installations 2.3 to 161 kV (7). 



LOW-VOLTAGE FUSES 

Plug fuses and cartridge fuses are the two principal 
categories of standard low-voltage fuses, and they are 
classified as non-time-delay, time-delay, dual-element, or 
current-limiting (13). There are also miscellaneous and 
nonstandard fuse classes. 

As with circuit breakers, there are three general fuse 
ratings (7); 

1. Current. The maximum dc or rms ac, in amperes, 
which the fuses will carry without exceeding a specified 
temperature rise limit (available range: milliamperes to 
6,000 A). 

2. Voltage. The maximum ac or dc voltage at which 
the fuse is designed to operate (usual low-voltage ratings 
are 600, 300, 250, or 125 V ac or dc or both). 

3. Interrupting. The assigned maximum short-circuit 
current that the fuse will safely interrupt (typical ratings 
are 10,000-, 50,000-, 100,000-, or 200,000-A symmetrical 
rms). 

Special ratings are also given to current-limiting fuses to 
specify the maximum current and energy the device will 
let through to the protected circuit when clearing a fault 
(7). 



236 



Plug fuses are rated at 125 V and are available with 
current ratings up to 30 A. Their use is thus limited to 
circuits with this voltage rating or less, except that they 
may be employed on systems having a grounded neutral 
where the maximum potential to ground of any conductor 
does not exceed 150 V (7). As a result, plug fuses have 
limited application in mine power systems (although an 
extensive popularity still exists for homes). Cartridge fuse 
applications, on the contrary, are widespread, to the point 
where mention of a fuse implies a cartridge. Figure 9.20 
shows the three standard low-voltage cartridge-type fuses 
(7). 

Non-Time-Delay Fuses 

As the name implies, these have no intentional built- 
in delay. They have a very simple construction, consisting 
of two end terminals joined together by a copper or zinc 
fusible element. The link is more current sensitive to 
melting than to time. Non-time-delay fuses are available 
as one-shot (or nonrenewable) and renewable; the former is 
the oldest cartridge fuse type in use today (7). With the 
one-shot, the link is in a sealed enclosure and the entire 
cartridge must be replaced after interruption. The renew- 
able fuse can be disassembled, and the link replaced. The 
lack of intentional time delay and a limited interrupting 
rating of around 10,000 A have substantially reduced the 
popularity of these fuses in recent years. 

Time-Delay Fuses 

The metal alloy used in time-delay fusible links is not 
only sensitive to current but also to the time period 
involved. In other words, a specific current existing for a 
specified time period is necessary to cause the heat- 
melting energy of the alloy. Such an arrangement permits 
harmless high-magnitude, short-duration currents to ex- 
ist, which are oftentimes necessary for proper system 
operation, as in motor starting. 

Dual-Element Fuse 

Originally designed primarily for motor-circuit pro- 
tection, the dual-element fuse (fig. 9.21) combines the 
features of non-time-delay and time-delay units. The time- 
delay or thermal cutout is provided for overload protection, 
while two fuse link elements give short-circuit protection, 
blowing in a fraction of a cycle on heavy currents. The 
thermal cutout will allow the passage of currents as high 
as five times its continuous rating for up to 10 s. Hence, 
these fuses may be matched closely to protect the actual 
motor running current and at the same time be sized to 
protect wiring and other equipment, and provide both 
these functions without nuisance blowing. In fact, prop- 
erly sized dual-element fuses are required on all fuse- 
protected trailing cables. They are available with up to a 
200,000-A symmetrical rms interrupting-current rating, 
and for further protection, most dual-element fuses also 
have a current-limiting feature. 

Current-Limiting Fuses 

Short-circuit protection requires that a fuse limit the 
energy delivered by the short circuit to a faulted compo- 
nent. Obviously, the energy any interrupting device lets 
through under fault conditions cannot exceed the pro- 



fl-rnn ^=2ZX^ 



Q) 

(8 



T¥T 



Ferrule type Knife-blade type 

0-60 A 70-600 A 

Figure 9.20.— Common cartridge fuses. 



<2> 



Bolt type 
601-6,000 A 



Multiple-bridge 
short-circuit link 



Fiber tube 




Quartz sand 
filler 



Alloy time-delay 
element 



Figure 9.21.— Inside view of dual-element fuse. 



tected components withstand rating. Current-limiting 
fuses provide this protection by restricting or cutting off 
fault currents before damaging peaks are reached. With 
very high fault currents, they are extremely fast, limiting 
current in less than one-quarter cycle, with current inter- 
ruption occurring within the first one-half cycle. Only a 
portion of the destructive short-circuit energy that is 
available is let through. By this, the current-limiting fuse 
allows the use of lower momentary and interrupting 
ratings by cutting off current within equipment ratings 
(7). Figure 9.22 illustrates how the fuse operates: the large 
waveform represents the available short-circuit current on 
a faulted system, and the performance of the fuse is 
superimposed. 

Restricting energy is a means of limiting the mechan- 
ical and thermal stress imposed on equipment that is 
carrying fault current. To illustrate this energy, consider 
figures 9.22 and 9.23 and the peak let-through current, L. 
It has been found that the magnetic forces during a fault 
vary as the square of fault current, I 2 , (7). These forces 
translate to mechanical stress, which could damage trans- 
former frames, bus structures, or cable supports. The 
let-through energy, I 2 t, represents a measure of the heat- 
ing effect or thermal energy of the fault with or without 
the fuse (with the fuse, the value is I 2 ,). I 2 t actually equals 
ji 2 dt, the time integral of the current squared for the time 
under consideration (8). Both I 2 , and I^t can be consider- 
ably reduced when current-limiting rases are used (7). 
Furthermore, equipment with an I 2 t withstand rating can 
be matched with the energy let-through limit of the fuse. 

Standard Fuses 

As implied by the foregoing, cartridge fuses come in a 
wide range of types, sizes, and ratings. Various classes for 



237 



Peak available current 





h— I 


— *» 






II 

I 






i 


/ '/ 2 cycle \ 


1 




A , 


jL ,*-— Arc \ 

r fy/x \ 



Peak let-through -fcH-Clearing 

current t j me 

Melting _| [_ Arcing 
time time 

Figure 9.22.— Current-limiting action of fuses. 



I 2 t 



Time 



rms 
current 




Figure 9.23.— Energy-limiting action of fuses. 



low- voltage units have been standardized (15), and a 
listing of general-purpose fuses follows (the first value 
listed is the range of continuous currents): 

Class G: to 60 A, 300 V to ground maximum, 
100,000-A symmetrical rms interrupting, current limit- 
ing, fit only class G fuse holders. 

Class H: to 600 A, 250 and 600 V, interrupting 
capacity up to 10,000 A, either one-time or renewable 
construction, commonly termed the "old NEC fuse." 

Class J: to 600 A, 600 V, 200,000-A symmetrical rms 
interrupting, current limiting, fit only a class J fuse 
holders. 

Class K: to 600 A, 250 and 600 V, 50,000-, 100,000-, 
or 200,000-A symmetrical rms interrupting, have the 
greatest current-limiting effect of all low-voltage fuses 
(available as straight current limiting, dual-element cur- 
rent limiting, and dual-element time-delay current limit- 
ing), fit class H fuse holders. 

Class L: 601 to 6,000 A, 600 V, 200,000-A symmetrical 
rms interrupting, current limiting, bolt-in mounting. 

Class R: to 600 A, 250 and 600 V, 200,000-A 
symmetrical rms interrupting, current limiting similar to 
class K level 5 fuse, fit only class R fuse holders. 

Class T: to 600 A, 250 and 600 V, 200,000-A 
symmetrical rms interrupting, current limiting but effect 
less than class J fuses, fit only class T fuse holders. 

Nonstandard Fuses 

Nonstandard fuses receive their name because of their 
special dimensions or use in special applications; they are 



not general-purpose fuses (7). Of the many available, four 
have important applications in mining: 

Cable Limiters. These fuses are for use in multicable 
circuits (paralleled cables) and are placed in series with 
each cable in parallel. They are designed to provide 
short-circuit protection to each cable, removing it from 
power in case of failure. Cable limiters are rated according 
to cable size (AWG 4/0 and so forth). 

Semiconductor Fuses. These devices are available in 
two types: semiconductor-protection fuses or semiconductor- 
isolation fuses. Both are used in series with the applica- 
tion. Protection fuses are employed where solid-state de- 
vices are to be protected rather than isolated after a 
failure; they have lower let-through characteristics than 
other current-limiting fuses. A specific application is pro- 
tecting a rectifier or thyristor in case of an overload 
current. Isolation types are high-speed fuses, used to 
isolate a defective solid-state device in case of its failure. 
These are mandatory fuses for individual power diodes 
paralleled in large rectifier banks. 

Capacitor Fuses. Capacitor fuses are applied in series 
with power-factor (pf) correction (or other type) capacitors 
and are used to isolate a failed component by clearing 
short-circuit current before excessive gas is generated in 
the capacitor. 

Welding Fuses. These are current-limiting fuses for 
use in welder circuits only. The time-current characteris- 
tics are such that these fuses allow a longer intermittent 
overload than general-purpose fuses, but still provide 
short-circuit protection. 



HIGH-VOLTAGE FUSES 

High-voltage fuses provide usable protection for 2.3- 
to 161-kV systems and fall into two general categories: 
distribution fuse cutouts and power fuses (7). Distribution 
fuse cutouts were designed for overhead distribution cir- 
cuits, such as the protection of residential distribution 
transformers. Even though their employment in utility- 
type systems is extensive, their use in mining is limited 
and in some cases restricted. Power fuses are another 
matter, as certain types offer extremely practical protec- 
tion in mine power systems. They can be applied to 
substation, distribution, and potential transformers (in 
series with the primary) and occasionally to distribution 
circuit conductors. For surface mine systems, the fuses are 
often equipped with contacts, arranged so that the fuse 
and its mounting act as a disconnect switch (fig. 9.24). 
There are two basic power fuses, expulsion and current- 
limiting types, and the next few paragraphs will discuss 
their operation, ratings, and application. 

Expulsion Types 

As with low-voltage fuses, high-voltage types start the 
current-interruption process by the melting of a fusible 
link, but as might be expected, deionization of the atten- 
dant arc becomes the most substantial part of current 
termination. To help the process, as shown in figure 9.25, 
the link is held under tension by a coil spring; upon 
melting, the spring pulls the contacts apart, lengthening 
the arc (4). In expulsion fuses, gases are liberated from the 
lining of the current-interrupting chamber by the heat 
generated from the arc. Both the earliest form of expulsion 



238 




attachment 



Figure 9.24.— High-voltage power fuse and support. (Courtesy 
S&C Electric Co.) 



Fiber tube -, 
Boric acid 
Plunger 




Strain element 
Main fuse 



Gap -, Disk 



Figure 9.26.— Cross section of boric acid power fuse refill. 



B- 




Fusible link Spring Glass tube 

/ * 




E 



Flexible lead 

Figure 9.25.— Fusible element under spring tension in high- 
voltage fuse. 



fuse and distribution fuse cutouts use a liner of organic 
material to deionize the generated gases by expelling 
them from the fuse holder tube to the surrounding air. The 
problem with this operation is the attendant flame expul- 
sion and loud noise. Hence, expulsion fuses are suitable 
only for outdoor usage, generally in substations remotely 
located from human habitation (7). 

The limited interrupting capacity (table 9.7) and 
unsuitability for indoor use of early expulsion fuses led to 
the development of the boric acid or solid-material fuse (7). 
Here, the interrupting chamber is made of solid boric acid. 
When exposed to arc heat, the material liberates steam, 
which can be readily condensed to liquid by venting the 
gas into a cooling device. The result is an operation with 
negligible or harmless flame and gas emissions and noise 
levels. The range of voltage, continuous current, and 
interrupting ratings is also greatly expanded. 

High-voltage boric acid fuses are manufactured in two 
styles (7): the fuse unit (nonrenewable), where the fusible 
unit, interrupting element, and operating element are all 
combined in an insulated tube; and the refill unit or 
fuseholder (renewable), where only the refill unit is re- 
placed after interruption. Figure 9.26 shows the internal 
components of a refill unit, while figure 9.27 illustrates 
the construction of the entire fuse. Table 9.7 provides a list 
of typical ratings for both styles. The fuse-unit style is 
intended for outdoor use at system voltages of 34.5 to 138 
kV, while the refill unit can be used indoors or outdoors on 
the surface at 2.4 to 34.5 kV. 






KEY 

A Fuseholder 

B Spring-and-cable assembly (copper cable carries 
load current ) 

C,D Fuseholder upper contacts and latch 

E Fuseholder lower contacts and latch 

F Refill unit 

Figure 9.27.— Disassembled refill unit for boric acid fuse. 
(Courtesy S&C Electric Co.) 



Current-Limiting High-Voltage Fuses 

High-voltage current-limiting or silver-sand fuses 
have the same advantages as previously discussed for 
low- voltage fuses and are of two different forms (7): those to 
be used with high-voltage motor starters for high-capacity 
distribution circuits at 2,400 and 4,160 V and those for use 
with potential, distribution, and small power transformers 
from 2.4 to 34.5 kV. The operation of either form is such 
that the arc established by the melting of the fusible 
element is subjected to mechanical restriction by a powder 
or sand filler surrounding the fusible element. The tech- 
nique provides three important features: 

• Current is interrupted quickly without arc-product 
or gas expulsion. This allows use indoors or in small-size 
enclosures on the surface or underground. There is no 
noise from the operation, and since there is no gas or flame 
discharge, only normal electrical clearances need by met. 



Table 9.7.— Ratings of high-voltage power fuses 



239 



.- , . fc , Boric acid fuse, Boric acid fuse, 

Nominal Expulsion-type fuse 1 -shot type renewable 

rating, Maximum Maximum Maximum Maximum Maximum Maximum 

kV continuous interrupting continuous interrupting continuous interrupting 

current, A rating, MVA 1 current, A rating, MVA 1 current, A rating, MVA 1 

2.4 — — — — 200,400,720 155 

4.16 — — — — 200,400,720 270 

4.8 — — — — — — 

7.2 100, 200, 300, 400 162 — — 200, 400, 720 325 

14.4 100,200,300,400 406 — — 200,400,720 620 

23 100,200,300,400 785 — — 200,300 750 

34.5 100,200,300,400 1,174 100,200,300 2,000 200,300 1,000 

46 100,200,300,400 1,988 100,200,300 2,500 — — 

69 100, 200, 300, 400 2,350 100, 200, 300 2,000 — — 

115 100,200 3,110 100,250 2,000 — — 

138 100,200 2,980 100,250 2,000 — — 

161 100, 200 3,480 — — — — 

1 3-phase symmetrical rating. 

NOTE.— Dashes indicate that standard fuses are not available in the specific voltage rating. 



Current-limiting 


fuse 




Maximum 


Maximum 


continuous 


interrupting 


current, A 


rating, MVA 1 


100, 200, 450 


155, 210, 360 


450 


360 


100, 200, 300, 400 


310 


100, 200 


620 


50, 100, 175,200 


780-2,950 


50, 100 


750-1,740 


40,80 


750-2,600 



• Very high interrupting ratings are available so 
these fuses can be applied on systems with very high 
short-circuit capacity (within their voltage rating). 

• All of the advantages of current-limiting action are 
available for high voltage. 

Table 9.7 provides a listing of typical ratings for 
current-limiting fuses. Instead of being rated by current, 
these fuses can also be "E-rated" (for instance, 100 E 
instead of 100 A), "C-rated," or "R-rated." The specifica- 
tions for E and C ratings are as follows: 

• E-rated fuses: 100 E and below, open in 300 s at an 
rms current within the range of 200% to 240% of the 
continuous rating of the fuse element; above 100 E, open 
in 600 s at an rms current within the range of 220% to 
264% of the continuous (or E) rating; 

• C-rated fuses: open in 1,000 s at an rms current 
within the range of 170% and 240% of the C ratings. 

E-rated fuses are considered as general-purpose or backup 
fuses, while R-rated devices are intended for use with 
high-voltage motor starters (7). 

Load-Break Switches 




Fused load-break switch 



It is possible that after the occurrence of a short 
circuit on a fuse-protected three-phase system, only one of 
the three fuses could open. Here current through the 
remaining two fuses might be reduced so that they do not 
open. The system then becomes single phased, which can 
cause serious damage to equipment. In a low-voltage 
circuit, dual-element fuses that are closely matched to the 
overcurrent point can usually handle the situation. On 
high-voltage systems, the problem is much more difficult 
when protection is by fuses alone. However, to take advan- 
tage of the lower cost of fuses and load-break switches 
versus the cost of a high-voltage circuit breaker, some 
manufacturers produce load-break switches with incorpo- 
rated high-voltage fuseholders. An example is shown in 
figure 9.28 where the fuses are interlocked to trip the 
operating mechanism of the switch if one or more of the 
fuses fail. Interlocking is usually accomplished with spe- 
cial high-voltage fuses that contain a spring-loaded 
plunger. Fuse activation releases the plunger, which trips 



h j 



J< 



I 
Switch 



-»-» 



XT\ 






Large fuse 



Actuator fuse 



Remote 
signaling 
circuit 
or switch- 
opening 
circuit 



Schematic showing interlocks 



Figure 9.28.— Load-break switch with interlocked high- 
voltage fuses. (Courtesy Line Power Manufacturing Corp.) 



240 



the switch mechanism. Precautions must be observed 
when using or considering these devices, and these are 
discussed in chapters 12 and 13. 

RELAYS 

Relays perform a major role in power-system protec- 
tion, where their purpose is to detect voltage and current 
anomalies. They normally receive information about sys- 
tem conditions through transformers or resistors, which 
reduce system parameters down to levels that the relays 
can handle. Upon detection of a problem, a relay operates 
to supply or remove control power to the shunt or UVR 
tripping elements of the switching apparatus. 

Because of their function, relays are sometimes called 
sensing devices. While transformers might also be consid- 
ered sensing devices, their function in protective relaying 
is solely as transducers. 

There are four basic relay types: thermal, electromag- 
netic attraction, electromagnetic induction, and static. 
(D'Arsonval movements are actually considered another 
relay type, but their operation is completely covered in 
chapter 5). The first three are electromechanical devices, 
and the following paragraphs will present their operation. 
Static or solid-state relays are discussed in chapter 12, 
because of related content. 

Relay Terminology and Types 

When a relay operates, it is said to close or open its 
contacts (9). Most relays are restrained by spring control 
and assume a specific position, either open or closed, when 
deenergized: hence there is a normally closed or NC 
contact and a normally open or NO contact. Symbols for 
both situations are shown in figure 9.29. 

When a relay operates to open NC contacts or close 
NO contacts, it is said to pick up the contacts, and the 
smallest actuating quantity to cause contact operation is 
referred to as the pickup value. When a relay operates to 
close NC contacts or open NO contacts, it is said to reset or 
drop out, and similarly, the largest actuating quantity to 
cause reset is the reset value of the relay. When the relay is 
deenergized to reset, the reset value is almost always 
greater than zero and is often specified as a percentage of 
normal operation. Most relays have adjustments or tap 
settings to adapt them to as wide an operating range as 
possible. 

The word describing relay operation has a formal 
meaning; for example, overvoltage relays, overcurrent re- 
lays, overtemperature relays, and so forth. Here the suffix 
refers to the actuating source (voltage, current, etc.), and 
the prefix "over" means that the relay picks up to close a 
set of NO contacts (or open NC contacts) when the actuat- 
ing quantity exceeds the magnitude at which the relay is 
adjusted to operate. Similarly, undervoltage, undercur- 
rent, and undertemperature relays reset to close NC 
contacts (or open NO contacts) when the actuating quan- 
tity decreases below a predetermined level. Some relays 
have both "over" and "under" functions (7, 10). 

Even with these definite meanings, common usage of 
relay terminology is rather straightforward. Pickup is 
used to refer to the point where the relay changes from its 
normal state to indicate a malfunction, while reset implies 
that the relay returns to its normal position. The normal 
position may occur when the relay is energized or deener- 
gized and depends on the application. 



Relays designed for protective circuits are usually 
provided with some means of visual indication that a 
specific relay has operated to trip a circuit breaker. These 
operation indicators or targets are often brightly colored 
and are operated mechanically or electrically. 

Specific relay types have been developed to meet 
special or general system-protection needs. Thermal relays 
serve directly or indirectly to measure power-system tem- 
peratures. Electromagnetic-attraction relays are used to 
instantaneously detect voltage and current changes. 
Electromagnetic-induction relays allow a time delay be- 
tween relay detection and contact action. Directional re- 
lays can sense the direction of current flow. 

Thermal Relays 

Thermal relays most commonly employ bimetallic- 
driven contacts with an operation similar to that described 
for the molded-case circuit breakers. Another approach is 
to use ambient temperature, as in the temperature- 
monitoring protector shown in figure 9.30. This is a sealed 
bimetallic thermostat that opens or closes at a specific 
temperature; it can be used, for example, to sense motor 
overtemperature if mounted against the end turns of a 
motor winding. 

Yet another bimetallic approach is to employ a heater 
element within the relay enclosures, connected in series 
with the circuit under consideration, as illustrated in 
figure 9.31A. The relay trip point for opening or closing 
the contacts is expressed in amperes, but is also a function 

Normally open Normally closed 

Figure 9.29.— Relay contact symbols. 



Insulation 




Device 




i^ss 



Bimetallic 
strip 




Figure 9.30.— Temperature-monitoring protector. 



Heater 



Thermal relay unit 





—•-To motor 



To 

• magnet 

coil 



Normal position 
A Bimetallic relay 



B Melting-alloy relay 



Figure 9.31.— Electromechanical-thermal relays. 



241 



of time and is determined by the heater rating. The trip 
setting is commonly based on a 40° C ambient tempera- 
ture, but the relay may be ambient or nonambient com- 
pensating. Most relays of this type must be manually reset 
after tripping. 

An electromechanical-thermal device not using bime- 
tallics is the melting-alloy or eutectic-alloy relay, figure 
9.31B. Being shock resistant and having high contact 
force, this is considered one of the most reliable thermal 
relays available, but because of its cost, it is not nearly as 
popular as the bimetallic type. The alloy melting point is 
extremely precise and is again related to a specific 
current-time characteristic. The relay can be reset after 
tripping and alloy resolidification. 

Two other thermal devices, resistance or thermistor 
types and thermocouples, operate with associated elec- 
tronic equipment to provide very precise temperature 
sensing and relaying. Here, for example, a probe can be 
inserted or embedded in a transformer or a motor winding 
to provide a spot temperature response. This type of device 
is very popular especially where large horsepower or 
capacity is involved. 



Electromagnetic-Attraction Relays 

There are three electromagnetic-attraction relays in 
common use: the solenoid, the clapper, and the polar (20). 
Although their operational speed might vary, all are 
considered instantaneous relays, since there is no built-in 
delay for pickup or reset. The solenoid and clapper types 
are available for ac or dc and are voltage or current 
actuated. Coil impedance is high for voltage and low for 
current. Polar units are dc sensing only, but may be used 
on ac circuits through rectification. All electromagnetic 
relays are available with NO contacts, NC contacts, or 
both. 

In solenoid units, the relay contact movement is 
initiated by a plunger being drawn into a cylindrical 
solenoid coil. Typical operating times are 5 to 50 ms, with 
the longer times associated with operation near the min- 
imum pickup value (20). A cross-sectional sketch of a 
solenoid relay is given in figure 9.32A. 

Four different clapper relays are shown in figure 
9.32B. These have a magnetic frame with a movable 
armature and operate by the attraction of the armature to 




Adjusting core screw 

Coil area 

I- Magnetic frame 

Nonmagnetic ring 

- Plunger 
Contacts 

Helical spring 



A Solenoid-type relay 



Mainspring — M\ 
Core- 




^f^r 



Magnetic frame 
Coil 

Residual pin 



Target - 

Armature Moving contact 
Indicating contact switch (ICS) 




^-Normally closed (break) contact 



"-Normally open (make) contact 



Contact multiplier 



Adjustable core - 



Target 



■-Magnetic frame M ^ in 9 contact 

i ji r ■■ Stationary__^f-~^n /Armature 

011 contact jpfer__ ^j|g~ ~ Armature adjusting screw 

jydfcLaTloop Insulating card- " 

p__^_ 52 Armature Residual plating 

Coil- 



Mi! J) Moving contact 
Lag loop ^zy 

Indicating instantaneous trip(IIT) 




Core 



^--Magnetic frame 



High speed 
B Clapper-type relay 
Figure 9.32.— Solenoid and clapper relays. (Courtesy Westinghouse Electric Corp.) 



242 



an electromagnetic pole {20). The armature controls the 
pickup or reset of contacts. 

As illustrated in figure 9.33, polar relays have a 
hinged armature in the center of the magnetic structure, 
which is here shown as an electromagnet but may be a 
permanent magnet. The relays operate when dc is applied 
to the actuating coil, and the polarity of the actuating 
source determines armature action, be it stationary or 
movement in either direction (10). In some units there is 
no retaining spring, and through a combination of con- 
tacts, the relays can sense actuating current through the 
coil in either direction (20). 

The pickup and reset values of clapper units are less 
precise than those of solenoid and polar relays; thus, 
clapper relays are used often as auxiliary or go, no-go 
devices (20). A common use for polar relays is in dc circuit 
protection where the actuating source is obtained from a 
shunt or directly from the circuit (10). 

A characteristic that should be considered when ap- 
plying any electromagnetic-attraction relay is the large 
difference that can exist between pickup and reset values. 
When an attraction relay picks up, the air gap is short- 
ened, and a smaller coil current is needed to retain pickup. 
Thus, the reset current may be much lower than the 
pickup current. The disparity is usually expressed as a 
percent ratio of reset current to pickup current, and is less 
pronounced in ac than dc relays. The ac relays can have a 
reset up to 90% or 95% of pickup, but dc ratios range from 
60% to 90% (10). This is no problem in overcurrent appli- 
cations where relay coil current drops to zero after pickup, 
but it is a concern where reset values are important. 

Electromagnetic-Induction Relays 



Contacts 



Stop 

o 



Control < 
spring < 




Actuating 
coiU 

umx >■■■ 



Stop 



Movable 
armature 

J Polarizing 

I COil v 



Pivot 



r 



Polarizing 
magnet 






To actuating quantity 

Figure 9.33.— Polar relay. 



Butt joint 
Main coil 




Lag coil 



Keeper - Disk air gap 

Figure 9.34.— Common induction-disk relay. 



Electromagnetic-induction relays are of two general 
types: induction disk and cylinder (20). Depending on the 
design, the induction-disk unit can be either a single- 
quantity or directional relay, whereas cylinder relays are 
intended to be directional. A single-quantity relay, as 
might be supposed, is actuated by and compares two 
sources (10). The most commonly used time-delay relays 
for system protection employ the induction-disk principle 
(7). 

Single Quantity 

Single-quantity time-delay relays of the induction- 
disk type use the same principle of operation that was 
described for induction motors in chapter 6, but the 
physical construction is quite different (20). A sketch of an 
elementary induction-type device is shown in figure 9.34, 
and most time-delay relays in use today have this arrange- 
ment. The disk, made of aluminum, is mounted on a 
rotating shaft restrained by a spring, and a moving 
contact is attached to the shaft (fig. 9.35). On one side of 
the disk is a three-pole electromagnet; the other side has a 
common permanent magnet or keeper. The operating 
torque on the disk is produced by the electromagnet, and 
the keeper provides a damping action or restraint after the 
disk starts to rotate. The retarding effect of the keeper 
creates the time delay or desired time characteristic of the 
relay. Figure 9.35 is a front-view illustration of an actual 
induction-disk relay removed from its drawout case; all 
important components are indicated. The unit pictured is 
for overcurrent, but overvoltage and undervoltage relays 
are also available and are identical in construction except 
for the electromagnet coil rating. 



nstonforssous unit 
CQiibrotion plate 




Control spring 



Crodfe 



Chassis contact 
block 



Figure 9.35.— Front view of induction-disk relay removed 
from case. (Courtesy General Electric Co.) 



The control spring carries current for the moving con- 
tact. If the actuating quantity driving the electromagnet is 
of sufficient magnitude and is sustained for enough time, the 
disk will rotate until the moving contact touches the station- 
ary contact. (Some relays use a lever on the moving disk that 
forces a pair of stationary contacts to close, so that no current 



243 



flows through the control spring and disk.) Pickup of these 
main contacts triggers the seal-in or time-delay element, 
which is an electromagnetic-attraction relay with its coil in 
series and contacts in parallel with the main contacts. When 
activated, this relay picks up and seals in, thus lightening 
the current-carrying duty of the main contacts as well as 
operating a target indicator. After pickup, it usually must be 
reset manually. 

The tap block at the top of figure 9.35 is to allow 
different tap settings on the electromagnet coil. Table 9.8 
lists the tap settings generally available in overcurrent 
relays (7), but some relays have wider ranges than those 
shown. Each range represents a different operating coil. 
Voltage relays have a narrower range of adjustment, 
because they are usually expected to operate within a 
limited change from the normal magnitude of the actuat- 
ing quantity (10). Be it a voltage or current relay, the coil 
and its tap settings are normally selected with respect to 
the ratios of the potential or current transformer used. 

Table 9.8.— Common current ratings of induction-disk 
overcurrent relays 

Time-delay elements Typical 

* instantaneous 

sl©m6nts 
Coil ran 9 e ' A Tap settings, 1 A adjustment range, A 

0.5 to 2.5 0.5, 0.6, 0.8, 1.0, 1.2, 1.5, 2.0, 0.5- 4.0 

2.5 2-16 

1.5 to 6.0 1.5,2,2.5,3.0,3.5,4.0,5.0,6.0 10 -80 

20 -160 

4.0 to 16 4.0, 5.0, 6.0, 7.0, 8.0, 10, 12, 16 (^) 

1 Tap settings will vary slightly according to manufacturer. 
Additional units are available for each time-delay range. 
3 Not adjustable. 



separate and relatively independent adjustment of the 
relay inverse-time characteristics. They are preset by the 
manufacturer, and the common responses are "inverse," 
"very inverse," "extremely inverse," "short time," and 
"long time," the first three being the most popular in 
mining. A comparison of these responses is given in figure 
9.37. The need for a specific response depends upon the 
application, and a few thoughts in terms of overcurrent 
relays follow (6). 

When the available fault-current magnitudes vary 
considerably, faster overall protection is usually gained 
with an inverse-time response. Very inverse curves provide 
the best overall protection where fault current remains 




MAGNITUDE OF ACTUATING 
QUANTITY 

Figure 9.36.— Inverse-time curve compared with definite- 
time curve. 



As shown in figure 9.35, overcurrent disk relays often 
have a second (auxiliary) ac-operated instantaneous ele- 
ment, which is a clapper-type relay (7). The unit is contin- 
uously adjustable over a calibrated range, and table 9.8 
lists some of these representative values. This relay oper- 
ates in series with the time-delay operating coil and is 
usually set to operate instantaneously at a current pickup 
value higher than that of the time-delay element. How- 
ever, since the same actuating source drives both ele- 
ments, the instantaneous-relay setting must be coordi- 
nated not only with the same source but also with the 
timed element. The instantaneous contacts can be in 
parallel with the time-delay contacts or can be connected 
to separate terminals. The unit also has a target indicator, 
which normally requires manual reset after tripping. 

The operational characteristic produced by the 
induction-disk principle is termed inverse time. Although 
mentioned earlier in this chapter, the inverse response is 
illustrated again in figure 9.36 to emphasize that the 
operating time becomes less as the magnitude of the 
actuating quantity is increased (10). The more pronounced 
this effect becomes, the more inverse the curve is said to 
be. All relay time curves are actually inverse, with the 
exception of a theoretical definite-time response. By defi- 
nition, definite-time characteristics imply that the operat- 
ing time of the relay is unaffected by the magnitude of 
actuating quantity. In reality, an actual definite-time 
curve is very slightly inverse (fig. 9.36). Regardless, the 
term definite time is normally applied to all fixed-time 
relays that approach this response. 

The control-spring tension, the damping magnet, and 
the magnetic plugs (A and B of figure 9.34) provide 



100.00 
60.00 

30.00 



10.00 
6.00 



UJ 

1 

1- 




Extremely 
inverse 



10 20 50 



MULTIPLES OF TAP 
VALUE CURRENT 



Figure 9.37.— Various time characteristics of induction 
units. 



244 



constant (detection of the fault, as seen by the relay, is 
mainly a function of fault location). Extremely inverse 
relays are designed to coordinate rather closely with power 
fuses and distribution cutouts and are also used in systems 
that have large inrush currents. The actual application of 
these characteristics in the mine is given in chapter 13. 

The operating time of an induction relay can usually 
be adjusted by selecting the distance of rotor travel from 
the reset to the pickup position (20). This is accomplished 
by adjusting the rest position of the moving-contact stop. 
The time dial, with evenly divided markings, facilitates 
positioning. When the response of the relay for different 
time dial settings is plotted, the result is a family of 
curves, an example of which is shown in figure 9.38. 
Current is plotted in terms of multiples of pickup, which 
enables the curves for a specific relay to be used with any 
tap setting. 

Directional 

The basic ac directional electromagnetic-induction 
relay or cylinder unit in common use is sketched in figure 

9.39. Its operation is similar to that of an induction motor 
that has salient poles for the stator, except that here the 
rotor iron is stationary and only the rotor conductor is free 
to rotate (10, 20). The rotor conductor is a thin- walled 
aluminum cylinder, and the two actuating quantities, 
causing I x and I 2 , independently produce torque on the 
cylinder. The cylinder drives a moving contact whose 
travel is restricted to a few degrees by the stationary 
contact and stops. Reset torque is established by a spiral 
spring. 

The ac directional relays are used to distinguish 
between current supplied in one direction or the other in 
an ac circuit, by recognizing phase-angle differences be- 
tween the two actuating quantities (10). (Conversely, a dc 
directional relay, or polar unit, recognizes differences in 
polarity.) To perform the ac comparison, one actuating 
value is used as a reference or polarizing quantity. There- 
fore, the polarizing quantity phase angle must remain 
fixed while the phase angle of the other fluctuates widely. 
One application of this technique is in power relays where 
the unit is polarized by circuit voltage, with circuit current 
being the other actuating value. Through this, the cylin- 
der detects power flow in one direction or the other. 
Another important application is an ac directional relay 
combined with an overcurrent relay, as shown in figure 

9.40. Here, tripping occurs only when the current has a 
specific relationship to the voltage, and power flow is in 
the tripping direction. 

BASIC RELAY CONNECTIONS 

In order to sense a malfunction and then supply 
tripping energy to the appropriate circuit breaker, a relay 
must be attached in some manner to the power system. 
Circuit connections for protective relaying are basically 
not too different from those discussed for instrumentation 
in chapter 5. Here, however, the relay coil receives the 
input information, and its contacts pick up or reset, thus 
affecting the control power to the circuit breaker. Direct 
relay connections to the monitored circuit are often re- 
stricted to low- voltage, low-power circuits because most 
relay current or voltage coils are designed to operate in the 
vicinity of 5 A or 120 V (4). Obviously, if power-system 
values exceed these levels, some interface is needed be- 



UJ 

»- 3 



3 


\ \ 




























Tie c 
ettir 

10 
K 


iia 

g 


1 












\ \ 


\ 


Ti 
\. s 








1 




















1 
i 






























.5 
.3 










_,, I "■-- 


££;;; 










.2 

1 














i/T 














1.5 2 3 4 


5 ( 


5 7 8 9 10 15 


20 







1.5 2 

MULTIPLE OF PICKUP 

Figure 9.38.— Family of inverse-time characteristics. 









t 1 


2 








1 . 








n <t> 








V 






. i 


2 

1 








j^ 






/ 




)V 




*, 






\ \ 1 


^ 


'I! 


,1 






, — 


JA^ 






r 






Cylinder 














• 


> 






Inner core 




@ 




@ 












ti 


2 






v F 


3 lug 



Laminations 



Figure 9.39.— Cylinder directional relay. 



tween the monitored circuit and the relays. Again, instru- 
ment transformers for ac and resistors for dc are used, a 
subject also introduced in chapter 5. 

There are five basic relay connections used for protec- 
tive relaying in the mining industry. For ac systems, these 
are direct, potential, and differential; and for dc work, 
direct and potential are used. Differential relaying is also 
available for dc, but the circuitry is not considered basic. 
Although some of the techniques are employed much more 
frequently than others, this section serves to introduce all 
these connections. 

Alternating Current Direct Relaying 

Direct relaying is used to sense the magnitude of 
current flow. As shown in figure 9.41 A, its simplest form 
consists of a current transformer (CT) secondary connected 
to a relay operating coil. Relay pickup current is thus a 



245 



function of line current. For instance, consider that the 
transformer ampere-turns ratio or current rating is 50/5 A 
or 10/1 and the relay pickup setting is at 0.5 A. This relay 
would theoretically pick up its contacts when line current 
is (10X0.5) or 5 A. The purpose of this connection is 
therefore to provide protective relaying for current in any 
conductor. 

The important items to consider in direct relaying are 
concerned with matching the performance of the CT with 
that of the relay. IEEE standards provide most of these (7). 

1. Ratios. As an obvious starting point after the 
foregoing example, standard ratios are listed below: 

Single-ratio CT, amperes: 



Coil terminal 



10/5 


200/5 


2,000/5 


15/5 


300/5 


3,000/5 


25/5 


400/5 


4,000/5 


40/5 


600/5 


5,000/5 


50/5 


800/5 


6,000/5 


75/5 


1,200/5 


8,000/5 


100/5 


1,500/5 


12,000/5 


3-ratio CT with centered-tapped secondary, amp 


25/50/5 




400/800/5 


50/100/5 




600/1,200/5 


100/200/5 




1,000/2,000/5 


200/400/5 




1,500/3,000/5 


300/600/5 




2,000/4,000/5 



Multiratio CT with multitapped secondary, amperes (cur- 
rent ratings higher than those shown are also available): 



600/5 



Rating 



1,200/5 



2,000/5 



Taps 
50/5 
100/5 
150/5 
200/5 
250/5 
300/5 
400/5 
450/5 
500/5 
600/5 

100/5 
200/5 
300/5 
400/5 
500/5 
600/5 
800/5 
900/5 
1,000/5 
1,200/5 

300/5 

400/5 

500/5 

800/5 

1,100/5 

1,200/5 

1,500/5 

1,600/5 

2,000/5 




Figure 9.40.— Directional overcurrent relay using induction- 
disk relay and cylinder relay. 



CT 



Line to be 
monitored 

/ 



L, H, 



Relay operating 
coil 

A Circuit connections 



'2 

x.4 



n 2 



■ • X? 



B Instantaneous current 



Figure 9.41.— Direct relaying in ac systems. 



The double-ratio and multiratio types provide flexibility 
through secondary taps. These values are for bushing-type 
or window-type CT's, which are the most popular in the 
industry. All these have the standard 5-A-rated secondary 
current. 

2. Secondary Current. The continuous-current rating 
of the secondary should be at least equal to the actual 
drain, but a full-load secondary current of 3 to 4 A is 
normal practice. An oversized CT is bad practice, as the 
percent error is much greater than with a correctly rated 
CT 

3. Short-Time Ratings. Both thermal and mechanical 
ratings should be considered. The thermal short-time 
value relates to the maximum symmetrical rms primary 



246 



current that the CT can carry for 1.0 s without exceeding 
its maximum specified winding temperature. The mechan- 
ical rating refers to the maximum asymmetrical rms 
current the CT can withstand without damage. In both 
cases, the rating is made with the secondary short- 
circuited. 

4. Voltage Rating. Standard voltage ratings are 600, 
2,500, 5,000, 8,700, and 15,000 V, and are the same as 
insulation classes found in mine systems. The CT will 
operate continuously at 10% above rated voltage without 
insulation failure. 

5. Burden. As defined in chapter 5, burden is the load 
connected to the CT secondary; expressions used are 
volt-amperes at a given power factor or an impedance with 
a power factor. The power factor is that of the burden. 
Table 9.9 lists standard values for CT's at 60 Hz. Relay 
burdens are so varied they cannot be listed, but chapter 10 
shows how CT burden and relay burden can be compared. 

6. Accuracy. Accuracy of a CT relates to its transfor- 
mation ability. In protective-relaying applications, accu- 
racy is not only important at normal circuit currents but 
also at fault-current levels. The problem in CT's is that 
core saturation leads to poor accuracy or ratio errors. 
Accuracy class designations use a C or T identifying letter 
followed by a classification number. C states that percent 
ratio error can be calculated, whereas T means that the 
value has been found by testing. The classification number 
relates to a standard secondary voltage of 10, 20, 50, 100, 
200, 400, or 800 V. At this voltage, the CT will deliver to 
a standard burden, 20 times normal secondary current 
with 10% ratio error or less, and it will not exceed 10% 
with any current from 1 to 20 times rated current with a 
lesser burden. (For example, C200 relates that for a 2.0-Q 
burden, (20X5) or 100 A can be delivered from the CT 
without exceeding 10% error. This error can also be 
calculated.) 

7. Polarity. Polarity relates to the correct phasing of 
primary and secondary currents, and figure 9.4 IB shows 
the relative instantaneous directions of current as per 
standard markings. This allows correct connections when 
more than one transformer is used, which is imperative in 
three-phase systems. 

As can be seen in the foregoing listings, actual man- 
ufacturer specifications should always be consulted before 
attempting to match CT's with relays for direct-relaying 
applications. 

Alternating Current Potential Relaying 

Potential relaying is as simple as direct relaying and 
enables circuit voltage to be monitored. Figures 9.42A and 
9.42S show two applications: sensing voltage across a 
resistance and between two conductors. There is a poten- 
tial transformer (PT) between the circuit and the relay. 
Figure 9.42C gives the polarity correspondence of instan- 
taneous voltages between the primary and secondary 
windings as well as conventional transformer markings. 
Standard PT's are single-phase, two-winding units con- 
structed so that the primary and secondary voltages 
always have a fixed relationship (7). 

To visualize the operation, consider figure 9.42A. The 
transformer is rated 2,400/120 V or has a 20/1 ratio, an 
overvoltage relay is used, and the relay coil is rated at 120 
V with the contacts set to pick up at 80% of rated. The 
contacts will therefore pick up when 1,920 V exists across 
the resistor. 



IEEE standards also provide guidelines for PT utili- 
zation, and a summary of these follows (7). In general, they 
are less rigorous than those for CT's. 

1. Voltage. Standard voltage ratios are available in 
table 9.10. When applied to sense voltage between two 
conductors, the nominal system voltage should be within 
+ 10% of the transformer nameplate rating. When used in 
three-phase mining systems supplying portable or mobile 
equipment, primary connections must be line to line. 
Special ratings, providing other than the standard 120-V 
secondary, are usually available. 



Table 9.9.— Standard burden for current transformers 

„ , . , ._.. Characteristics for 60 Hz and 

Standard Genera characteristics _ . . , 2 

oianaara 5.^ secondary current 2 

burden 

designation 1 Resistance Inductance Impedance Apparent power pf 
(R), fi (L), mH (Z), fi (S), VA 

B-0.1 0.09 0.116 0.1 2.5 0.9 

B-0.2 .18 .232 .2 5.0 .9 

B-0.5 45 .580 .5 12.5 .9 

B-1 .5 2.3 1.0 25 .5 

B-2 1.0 4.6 2.0 50 .5 

B-3 2.0 9.2 4.0 100 .5 

B-4 4.0 1^4 aO 200 .5 

1 B-0.1, B-0.2, and B-0.5 are usually applied for metering purposes; B-1 
through B-4 are usually applied for relaying. 

2 At 5 A, S = l 2 2; for example, for B-2, S = 5 2 2 = 50 VA. 



Table 9.10.— Standard ratings for potential transformers 



Primary, V 
120 


(Seconds 
Ratio 

1/1 

2/1 

4/1 

5/1 
20/1 
35/1 
40/1 
60/1 


iry, 120 V) 

Primary, V 
8,400 


Ratio 
70/1 


240 


12,000 


100/1 


480 


14,400 


120/1 


600 


24,000 


200/1 


2,400 .. 


36,000 


300/1 


4 200 


48,000 


400/1 


4 800 


72,000 


600/1 


7,200 







•Line monitored 



, Relay 
coil 




Lines 
monitored 



A Resistance 



B Between conductors 



Instantaneous 
direction of ■<T 
current flow ^ 




Relation 

between 

X 4 voltages 



C Polarity 
Figure 9.42.— Potential-relaying connections. 



247 



2. Accuracy. Ratio and phase-angle errors of standard 
PT's are usually so small they can be neglected, and any 
standard transformer is satisfactory as long as it is used 
within its thermal and voltage limits. If the transformer 
load is within rated burden, the transformer is suitable 
over the range from zero to 110% of rated voltage. Regard- 
less, standard accuracy classes do exist for PT's, ranging 
from 0.3 to 1.2. These values represent the percent ratio 
corrections to obtain true ratio. 

3. Burden. The burden of a PT, or thermal burden 
limit, is expressed in voltamperes. It is usually sufficient 
to add the voltampere ratings of parallel loads arithmeti- 
cally to obtain a total voltampere burden. Accuracy is 
usually satisfactory at burdens well below rated, but the 
transformer voltampere rating should not be exceeded. 

4. Fusing. In some instances, fusing the primary of a 
PT is not advisable, especially when the protective- 
circuitry's function is to sense a critical overvoltage con- 
dition (for instance, monitoring the voltage across a 
grounding resistor). Yet when the PT is connected line to 
line, it must be protected in case of PT failure or secondary 
conditions that will lead to failure. General practice is to 
use current-limiting fuses, sized to the transformer full- 
load rating and installed in the primary circuit between 
each ungrounded conductor of the system. Fuses are 
preferred over circuit breaker primary protection because 
the latter is accessible for manual tripping. A major use 
for PT's in mine systems is to supply control power to 
protective circuitry; secondary protection in this case is 
unnecessary. For other loads such as branch circuits for 
120-V convenience outlets, the additional branches should 
be fused or protected by molded-case circuit breakers, the 
latter being general practice. 

Alternating Current Differential Relaying 

In differential relaying, a relay is operated by the 
vector difference of two or more actuating quantities, and 
relay pickup is determined by a difference threshold (10). 
Most applications of this scheme are of the current differ- 
ential type. A basic circuit is shown in figure 9.43A, where 
the dashed portion represents the area to be protected. 
Two matched CT's are interconnected, and an overcurrent 
relay is inserted between them. Under normal conditions, 
or even when a fault occurs outside the protected zone, the 
CT secondary currents will circulate and not flow through 
the relay coil. However, if the current in both CT primaries 
becomes unbalanced, current will flow through the relay 
in proportion to the vector difference of the current enter- 
ing and leaving the protected circuit (fig 9.43S). 

A problem with this basic circuit is that CT's are very 
difficult to match; on identical units, the same primary 
current will not always give the same secondary current 
(7). Thus, the relay must be set so that it does not pick up 
on maximum error current between the CT's. An approach 
that usually overcomes the mismatch problem is the per- 
centage differential connection. As illustrated in figure 
9.43 C, the main change is that the relay is now an over- 
current current-balance type (10). The differential current 
required to operate the relay is a variable quantity 
because of the relay restraining coil, and it offsets errors in 
the actuating sources. 

Direct Current Connections 

In addition to their popularity with ac systems, direct 
relaying and potential relaying are also the two most used 



protective relaying connections for dc systems. Direct 
relaying (fig. 9.44) consists of a dc overcurrent relay 
connected to a resistance shunt. The relay operating coil is 
matched to the shunt voltage at the desired pickup level 
(shunt full-load current rating usually gives 50 or 100 mV 
across the shunt). For low-current applications such as 
sensing dc current in a grounding conductor, current-relay 
operating coils are sometimes inserted in-line with the 
monitored conductor. Potential relays are also directly 
connected with the coils between the conductors of inter- 
est. Resistive dividers are at times employed to drop the dc 
system voltage down to the coil rating (as discussed in 
chapter 5). 




System portion 
to be protected 



External load 
or fault 



A Basic circuit under normal conditions 



A ,r 



2 



i 

Internal fault 



" it 



B Basic circuit under abnormal conditions 



Restraining 
coil 




C Percentage differential relay connections 
Figure 9.43.— Differential-relaying connections. 



Monitored 
line 
Shunt / 



Overcurrent 
relay coil 




Potential 
relay coil 



?> 



Voltage- drop 
resistor 



Direct relaying 



Potential relaying 



Figure 9.44.— Dc direct-relaying connections. 



248 



KINDS OF PROTECTION 

Several relaying terms describe the protection re- 
quired in many mine power systems: 

1. Undervoltage, 

2. Overload (sometimes called overcurrent), 

3. Short circuit, 

4. Ground overcurrent (or ground fault), 

5. Ground continuity, and 

6. Overtemperature. 

Classifications such as these are known formally as kinds 
of protection. The first five are necessary protection on all 
portable and mobile mining equipment, although excep- 
tions are provided within Federal regulations (17). This 
section expands the basic relaying material by describing 
how each kind of protection is used in the mine power 
system. The content is mainly pointed at high-voltage, 
three-phase ac mining systems and in general is restricted 
to relaying external from circuit breakers. Accordingly, 
these kinds of protection imply the following parameters: 

• Line-to-line voltages for undervoltage, 

• Line overcurrent for overload, 

• Three-phase or line-to-line faults for short circuit, 

• Faults causing zero-sequence current for ground 
overcurrent, and 

• Grounding-conductor resistance for ground continuity. 

Even though overtemperature is listed as item 6, it is 
usually applied to protect a specific component; thus, it 
will be discussed in chapters 12 and 13. 

Control Wiring 

Figures 9.45 and 9.46 show simplified diagrams of 
typical control wiring interconnections among the power 
source, relay contacts, and circuit breaker tripping ele- 
ments. In both diagrams, a potential transformer supplies 
120 Vac with its fused primary connected line to line. 

Figure 9.45 illustrates cases where the tripping ele- 
ment is a UVR. The contacts can either reset to remove 
power from the coil (contacts in series with coil) or close to 
short it out (contacts parallel the coil). In the latter case, it 
can be seen that a resistance is placed in series with the 
contacts. In fact, the UVR itself will trip the breaker if 
control voltage is decreased in the range of 40% to 60%. 

Basic shunt-tripping connections are given in figure 
9.46, where power is supplied to the element to cause 
tripping. Here the contacts for the various protective 
relays are paralleled, and the combination is in series with 
the trip coil; closure of any contact trips the breaker. It 
should be obvious in figure 9.46A that the power causing 
tripping is ac, while in figure 9.46B, it is dc. The capacitor 
in the dc circuit is employed for energy storage to augment 
tripping if there is a drop in the PT primary voltage when 
a relay contact closes. 

Phase Protection 

Phase protection by protective relaying can be over- 
load, short circuit, or both, depending upon the relays 
used. (Molded-case circuit breakers afford this same flexi- 
bility depending upon the internal tripping element used.) 
Time-delay relays are employed for overload, with instan- 
taneous units for short circuit. Figure 9.47 illustrates the 
combined protection for three line conductors, using an 




Power to 
load 



Relay contacts 
open to trip 

Removing power to UVR 



Fuses 




Power to 
load 



I 
Limit 
PT resistor ' UVR 

w» — t — *~0 — I 



T 
X 



Relay contacts 
close to trip 



Shorting out UVR 
Figure 9.45.— Typical control wiring for UVR. 



Line-to- 
line Manual 

U * ip 



Line-to- 

Ime 



PT 




Manual 
trip 

H f f oo— » 



. Shunt 
O trip 
coil 




N Shunt 
O trip 
coil 



A Simple 



B Capactor tripping 



Figure 9.46.— Typical control wiring for shunt-tripping ele- 
ment. 



Time-trip contact 



Instantaneous 
trip contact 




To trip 
circuit 



• Instantaneous-element current coil 
Time-element current coil 

1 Protected equipment 

Figure 9.47.— Three-phase overcurrent and short-circuit con- 
nections. 



249 



induction-disk relay (7). The three current transformers 
are placed in wye, driving the wye-connected operating 
coils. The time-delay element is set on as low a tap setting 
as practical, enabling protection for sustained moderate 
overloads. The instantaneous units, however, are set to 
pick up on a current value slightly higher than the 
maximum peak load, thereby affording protection against 
short circuits or enormous overloads. 

The device numbers that were presented in chapter 4 
are used extensively to describe the relay function. The 
number 51 signifies time-delay relays for ac overcurrent, 
and 50 is used for instantaneous devices. A combination 
instantaneous and time-delay ac overcurrent relay is often 
noted by 50/51. 

If the connections are as shown in figure 9.47 and 
transformer phase-angle errors are ignored, the secondary 
currents of each CT are in phase with the primary cur- 
rents, and each relay responds to abnormal conditions for 
its respective line (10). This also applies to figure 9.48A 
where the 50 elements are omitted. If line currents are 
approximately balanced, short-circuit protection for all 
three lines can also be provided with an open-delta con- 
nection, as in figure 9.485 (10). As might be expected, this 
approach is not as precise as the straight wye connection, 
but a third overcurrent relay may be inserted in the CT 
common connection for backup protection (see chapter 5 
for a similar discussion on instrumentation). An advan- 
tage of the wye connection that is lost when the open-delta 
approach is use is the ability to sense zero-sequence 
currents through residual relaying. Thus, two current 
transformers are rarely applied as the only means of 
circuit protection. 

Ground Overcurrent 

To this point, the chapter has basically considered 
power-conductor protective relaying, and the extremely 
important subject of ground-fault protection has received 
only terse reference. Various relay configurations may be 
utilized to provide ground-overcurrent protection, some of 
which are quite elaborate. However, nearly all these 
techniques fall into one of five broad classifications (7): 
direct relaying, potential relaying, residual connection, 
zero sequence, and broken delta. Direct relaying, potential 
relaying and zero sequence are frequently used in 
resistance-grounded mine power systems, with zero- 
sequence relaying being the most popular. Unless other- 
wise noted, the following discussion will assume that the 
system is resistance grounded. 

The point of application for direct and potential 
ground-fault relaying is usually restricted to the system 
neutral point or grounding resistor, whereas the other 
three techniques can provide protection anywhere in the 
system. Usually a combination is needed for complete 
assurance of clearing all ground faults. 

Direct Relaying 

The simplest form of ac ground-fault protection is 
direct or neutral relaying. A current transformer is placed 
about the grounding conductor and located between the 
neutral point of the source transformer and the grounding 
resistor, as shown in figure 9.49. The grounding conductor 
acts as the primary winding of the CT, while the secondary 
winding is connected to the ground-overcurrent relay 
(5 IN). If the current through the grounding conductor 




Line 
A 



Line 
B 



Line 
C 



H 



:r 



>^ 



A Wye connected B Open-delta connected 

Figure 9.48.— Two CT approaches. 



^Pb 




Neutral 

grounding 

resistor 



Grounding conductor 



Figure 9.49.— Neutral-resistor current-relaying scheme. 



exceeds a predetermined value, the relay acts to trip the 
circuit breaker. 

In many situations, some ground-current flow is nor- 
mal, due to system unbalance, capacitive-charging cur- 
rents, or inductive-coupling effects, and so the circuitry 
must be adjusted to pick up only when the normal level is 
exceeded. As will be seen, the pickup point should always 
be less than the system current level. 

The major disadvantage with this direct relaying 
method is that, should the grounding resistor or the 
grounding conductors become open, it will never detect 
any ground-current flow. The system will continue to 
operate with no abnormal indication, and then the system 
can become essentially ungrounded, posing a personnel 
hazard especially where resistance grounding is manda- 
tory (12, 14, 18). Accordingly, although the technique does 
find application on some portions of the ground system, 
some States do not allow its use on substation grounding 
resistors, even for a second line of defense. 

Potential Relaying 

Potential relaying, as shown in figure 9.50, is often 
used as a sole means of ground-fault protection at the 
surface substation and can also be used as a backup to 
other protection schemes at a unit substation or power 
center. With this method, the primary winding of a PT is 
connected across the neutral grounding resistor, while the 



250 



secondary winding is connected to a voltage-sensing 
ground-trip relay (59G). If current flows through the 
grounding conductor, a voltage is developed across the 
grounding resistor. When the voltage rises above a preset 
level, the ground-trip relay causes the circuit breaker to 
trip. 

Unlike direct relaying, potential relaying has the 
advantage of being able to detect a ground fault with the 
neutral grounding resistor in an open mode of failure. 
However, if the grounding resistor fails in a shorted mode, 
potential relaying is rendered inoperable. 

Zero-Sequence Relay 

Zero-sequence relaying, also termed balance-flux re- 
laying, is the most reliable first defense against ground 
faults in mine power systems. As shown in figure 9.51, the 
circuitry consists of a single window-type CT; the three 
line conductors are passed through the transformer core, 
forming the CT primary (6). On four-wire systems that 
have line-to-neutral loading, the neutral conductor must 
pass through the window but the grounding conductor 
must not. However, such loading is not allowed in mining, 
so only the line conductors can be used. 

In a symmetrical phase set, the vector sum of the 
three currents in the primary circuit will be zero, and no 
current will flow in the secondary. During a line-to-neutral 
fault, the phase unbalance will induce a current flow in 
the CT secondary, proportional to zero-sequence current. If 
the secondary current exceeds the relay pickup setting, 
circuit breaker tripping will be initiated. 

This phenomenon can be easily demonstrated. In 
terms of symmetrical components, the three phase cur- 
rents can be written as 

la = ^al + *a2 + ^aO> 
lb = Ibl + lb2 + IbO> 
Ic = Icl + Ic2 + IcO- 

The primary current, L rim , for the CT can be considered as 
the vector sum of the three line currents, which is also the 
current flowing through the grounding conductor (or that 
external to the CT window). Therefore, 



I P 



x. = la + lb + Ic 



= (a 2 + a + l)I al + (a 2 + a + 1)1^ + 3I a0 . 
Because a 2 + a + 1 = 0, 

Iprim = 3I a0 - 

Thus, for an unfaulted or balanced condition, 

Iprim = 3I a0 = 3I b0 = 3I c0 = 0, 

and no current is induced in the CT secondary. However, 
for a line-to-neutral fault involving phase a, the primary 
current equals the ground-fault current, If, or 

Iprim = If = 3I a0 , 

and a current is induced in the CT secondary to initiate 
tripping. 

Zero-sequence relaying is not affected by CT error, and 
therefore gives very sensitive tripping. The scheme is 
widely applied in mining at all voltages. 



Residual Relaying 

Residual relaying (fig. 9.52) is used in conjunction 
with CT's placed about the phase conductors. This relay- 
ing technique is used primarily on high-voltage distribu- 
tion circuits that require CT's and inverse-time relays for 
overcurrent protection. The CT's and the phase- 
overcurrent relays are both connected in a wye configura- 
tion. The ground-fault or residual relay is connected be- 
tween the neutral points of the CT's and the relays. 




PT and voltage relay (59G) 



Grounding conductor 



Figure 9.50.— Neutral-resistor potential-relaying scheme. 



Ckt 
bkr 



;^Rb 



;^^ 



:^Pf 



'ih 



Phase 
relays 




Zero-sequence 
relay (51G) 



Figure 9.51.— Zero-sequence ground relay connections. 



Ckt 
bkr 



P^Pt 



;=^R 



■ih 



Phase relays 
J 



L@J 



Residual relay (51G) 
Figure 9.52.— Ground relay in residual connection. 



251 



As the current flowing through the residual relay is in 
proportion to the sum of the line current, the principle of 
operation of the residual method is similar to that of 
zero-sequence relaying. However, because of errors due to 
CT saturation and unmatched characteristics, residually 
connected relays are often subjected to nuisance tripping. 
Hence, they cannot have sensitive or low pickup settings. 
This arrangement will not always provide consistent re- 
petitive tripping at the required tripping levels for mine 
power systems. 



Grounding 
conductor 




Broken-Delta Relaying 

Broken-delta ground-fault protection is somewhat 
similar to the residual method, except that the three CT's 
are wired in series, as shown in figure 9.53. The resulting 
output voltages from the transformers form a closed delta 
if the load is balanced (fig. 9.53A). An unbalanced condi- 
tion, such as a line-to-neutral fault, will cause the forma- 
tion of an open delta (fig. 9.535), and the resulting voltage 
causes current through the relay operating coil. The 
broken delta is sensitive to any unbalance, but the zero- 
sequence relay operates only on faults causing ground- 
current flow (6). 

Ground-Check Monitoring 

The effectiveness of all the ground-relaying methods 
depends upon the integrity of the grounding system. A 
ground-check monitor is the device used to continuously 
monitor the grounding connections to verify continuity 
(2-3, 9). If conductivity is inadequate, it is the function of 
the monitor to trip the circuit breaker that feeds power to 
the system experiencing defective grounding. 

As shown in chapter 7, the grounding conductor is not 
essential for mining machine operation, but it is impera- 
tive for personnel safety. The ground-check monitor en- 
hances safety by making sure, via the ground connections, 
that the equipment frames are at near-neutral potential. 
Again, the maximum allowable frame potential to earth is 
40 V on low and medium voltage and 100 V for high- 
voltage systems. 

Ground-check monitoring is an extensive subject, 
which can only be outlined here; the references listed at 
the end of the chapter can be consulted for more detail, 
particularly references 2-3, 9, and 11 . 

Although there are potentially numerous ways of 
monitoring ground continuity, only a few are considered 
practical to construct or have the required high reliability 
(9). These techniques can be divided into two general 
classifications: pilot monitors and pilotless monitors. Mon- 
itors in the mining industry use these techniques but are 
also referred to by different names, which will be discussed 
later. 

Pilot 

Pilot monitors use a pilot or ground-check conductor 
(see chapter 8) to perform the task and are of three general 
kinds: series loop, transmitter loop, and bridge (9). 

In the most common series loop circuit, a power 
supply, the relay operating coil (instantaneous contacts or 
a minimal time delay), the pilot conductor, and grounding 
conductors are connected as shown in figure 9.54 (9). If 
the pilot or grounding conductor breaks the loop or if the 
power supply fails, the relay contacts will reset. The 
circuit can be either ac, using 60-Hz line frequency, or dc. 



A Normal operations 



IbA U Ia "Ib w I 



'b+g 



firam 




v b 
v=o 



B Ground fault 
Figure 9.53.— Broken-delta protection. 



To 
circuit 
breaker 

trip 




Pilot 




•* 






Machine 
frame 






Voltage 
source 






Ground 


\ 


J Relay 






1 

To power-center 
frame ground 









Figure 9.54.— Series loop ground-check monitor. 



252 



The transmitter loop concept is basically the same as 
the series loop, except the voltage source is installed in the 
machine (fig. 9.55) (9). Here, the source must receive its 
power from the machine and the relay cannot pick up until 
the circuit is energized; therefore, the monitor must be 
temporarily bypassed in order to close the circuit breaker. 

Bridge-type monitors use the series combination of 
the pilot and grounding conductors as one leg of a Wheat- 
stone bridge. Figure 9.56 shows a general circuit, where Z 3 
is used to balance the bridge for a specific pilot and 
grounding-conductor impedance (9). Bridge output is 
sometimes amplified, but with or without amplification, 
the relay resets if a preset impedance level is exceeded. 
Bridge input can be 60-Hz ac, dc, or an audio frequency 
such as 5,000, 2,500, or 900 Hz (2, 9). 

Pilotless 

Pilotless monitors, as the name implies, do not use the 
pilot conductor. Instead, as shown in figure 9.57, an audio 
signal is placed on the phase conductors through a filter 
and removed at the machine through filters, completing 
its path back to the source in the grounding conductor (9). 
Instead of the filters, some models use coils similar to CT's 
to send and receive the audio signal. Between the ground- 
ing conductor and the power-center frame is a saturable 
reactor, which shows high impedance to the monitoring 
frequency. Its purpose is to restrict the monitoring signal 
to the intended path. This presents problems in coupler 
grounding, as the coupler metallic shell is commonly 
grounded to the grounding conductor as well as physically 
connected through its receptacle to the power-center 
frame. The grounding conductor must be isolated from the 
shell ground so that the reactor will not be bypassed. 

Problems and Requirements 

All of the basic techniques are plagued by some 
disadvantages, and the attempt to achieve a reliable 
monitor has been a perplexing experience for the mining 
industry. The reason is tied to the basic character of the 
mine power system: figure 9.58 provides a conceptual view 
of some difficulties that can arise (2). One of the more 
pronounced problems is the parallel ground paths estab- 
lished by contact through the mine floor or through 
grounding conductors on other machines. The alternate 
ground paths may have a resistance as low as the ground- 
ing conductor, but in the majority of cases these are very 
temporary in nature and thus cannot be relied upon. Stray 
ac and dc and induced ac are an ever-present problem in 
many mines. With dc rail haulage, for example, substan- 
tial direct current can stray from the rail when a poor bond 
is present and end up flowing in the ac ground system. If 
the cable has a G-GC or SHD-GC configuration (chapter 
8) or the system current is unsymmetrical, ac can be 
induced in the grounding conductors. At times, this cur- 
rent can be of significant magnitude, not only on under- 
ground loads but especially on surface excavating machin- 
ery. Another problem results from trailing-cable 
deterioration; in a splice, for instance, there is a chance 
the ground-check conductor could short to the grounding 
conductors. Power-system transients, occurring from light- 
ning or switching surges and wiper contacts on reeled 
units, present additional problems, and all of these situa- 
tions can affect ground-check monitors. 

MSHA has established several guidelines for low- 
voltage and medium-voltage monitors that must be met 



before a monitoring device is approved. In these regula- 
tions, two monitors are recognized: a continuity type and 
an impedance type. A continuity monitor is one that meets 
the general requirements of 30 CFR 75.902 (17). It moni- 
tors only the grounding-conductor continuity and does not 
measure impedance; pilotless techniques fall into this 



To 

circuit breaker 

trip 



Machine frame 



r\ 


Pilot 


J7^ 






Relay 


Ground 




1 








Transmitter 




I 





















To power-center 
frame ground 

Figure 9.55.— Transmitter loop ground-check monitor. 



To circuit breaker trip 



t-n-» 



Relay 





Preset level 
















Voltage 
source 




Pilot 




L/^i v\ 




Machine 
frame 

> 


V^3 


Ground 












Figure 9.56.— Bridge-type ground-check monitor. 



Line conductors 




Ground 




Saturable 
reactor 



To power-center 
frame ground 



Figure 9.57.— Pilotless ground-check monitor. 



253 



Machine 



Ground 




Grounding conductor 



Pilot conductor 



Cable -v 
coupler i 



Possible X 
fault 



Grounding conductor 



Safety § Earth contact % Earth contact 

ground bed 



Wiper contact if ree 



Earth Earth 

Figure 9.58.— Some difficulties associated with ground-check monitoring in mining. 




class. An impedance monitor requires a pilot conductor 
and monitors any change in the impedance of the loop 
formed by the pilot and grounding conductors. The bridge 
technique is therefore an impedance monitor. The relevant 
requirements for both monitor types are as follows (2, 9, 
U): 

1. The monitor must be "fail-safe"; in other words, the 
failure of any component, other than the relay contacts, 
must make the trip-circuit contacts reset. (The relay must 
pick up its contacts when in normal operation.) 

2. The monitor must not trip when (a) input voltage is 
varied by + 15% or - 20%, or (b) 5.0 V minimum to 25 V 
maximum, 60 Hz, or 10 A dc is introduced in the grounding 
circuit. These conditions are intended to ensure that the unit 
stays operational, even when under the influence of power- 
line fluctuations, stray currents, or induced currents. 

3. The open-circuit monitor voltage cannot exceed 40 
V rms. 

4. When detecting grounding-conductor continuity, (a) 
continuity monitors must trip the circuit breaker if the 
grounding conductor is broken at any point regardless of 
low-impedance parallel paths (75 Q is considered an open 
connection), and (b) impedance monitors must trip the 
circuit breaker if the impedance of the grounding circuit, 
external to the grounding resistor, increases to cause a 
40-V drop under fault conditions (or, by Ohm's law, 1.6 Q 
for a 25-A limit). 

5. Filters must not cause a personnel hazard during 
normal operation or when the grounding conductor is 
opened. 

6. The maximum time delay for contact reset, after an 
inadequate ground is detected, cannot exceed 250 ms. 

7. When two or more monitors are operated in paral- 
lel, no interference can occur to cause incorrect tripping. 

At this writing, similar guidelines for high-voltage 
ground-check monitors have yet to be established. How- 
ever, by 30 CFR 75 and 77 (17), the maximum open-circuit 
monitor voltage is established at 96 V rms. Furthermore, 
continuity monitors must adhere to item 4a above, with 
impedance monitors conceivably tripping if the grounding 
circuit impedance causes a 100-V drop. 



Advantages and Disadvantages 

As mentioned, each basic ground-check technique has 
inherent advantages and disadvantages. A listing of these 
is rather informative (2-3, 9). 

Other than simplicity, the advantages of series loop 
circuits are minimal. Designs using a dc source are im- 
mune to stray ac but not dc. Further, ac monitoring can be 
subject to nuisance tripping by stray dc current that offsets 
the signal current. However, when the relay operating coil 
is isolated by a blocking capacitor, an immunity to stray dc 
is gained. In any case, the relay coil must have a very low 
impedance to be sensitive to grounding-conductor imped- 
ance. Two disadvantages of the circuit are substantial: 
Parallel paths and grounded pilot conductors easily negate 
its operation. 

The advantages and disadvantages of the transmitter 
loop technique are basically the same as for series loop, 
except the circuit can detect pilot-to-ground shorts. 

Simple bridge monitors have the same problems as 
series loop models, but they are very sensitive to changes 
in grounding-conductor and pilot-conductor impedance. 
The more elaborate designs can be made immune to ac and 
dc stray currents, yet even these sometimes cannot distin- 
guish between a sound grounding conductor and an illegal 
parallel path. 

In general, pilotless monitors are superior to pilot 
designs, except they are obviously more expensive. Be- 
cause the ground-check conductor is not needed, all asso- 
ciated problems are removed. The most elaborate models 
can distinguish parallel paths and are immune to stray 
currents; however, the simple designs are vulnerable to 
both. Some pilotless designs also have the advantage of 
being adaptable to pilotless or pilot use, depending on the 
need. 

Pilot-conductor ground-check monitors can serve the 
very important function of safety interlocking. This fea- 
ture is required on many portions of the mine power 
system and is used on almost all high-voltage systems. An 
example of interlocking is shown in figure 9.59. The loop 
circuit sensed by the monitor not only includes the pilot 
and grounding conductors, but can also involve a series of 
contacts and switches. Whenever one of these is opened, 



254 



Upstream I 
power I 
equipment 



Top and side 
cover interlocks 




Figure 9.59.— Pilot interlocking circuit using ground-check 
monitor. 



the continuity of the ground-check circuit is compromised, 
the monitor resets its rely, and the circuit breaker trips the 
power to that section. Safety devices incorporated in the 
interlocks include top and side cover switches, which open 
if a cover is removed, emergency-stop switches, the pre- 
break contacts on couplers, and pilot-break monitors on 
disconnect switches. 

When such sophisticated interlocking is required, 
pilot-type monitors must almost invariably be used. For 
instance, pilotless models cannot be employed when in- 
line cable couplers are on the circuit. When there are 
gear-mounted couplers on the power equipment in the 
monitored circuit, pilotless monitors may be used if they 
are wired into the ground-check and grounding contacts of 
the couplers, so that they will trip the circuit breaker prior 
to plug removal. This applies to almost all low-voltage and 
medium-voltage monitored circuits in mine distribution. 



ARRANGEMENTS FOR MINING 

There are two groups of protective-relaying equipment 
within a typical power system: primary and backup. 
Primary relaying has the goal of clearing all faults and 
overloads, and aims to isolate offending power-system 
segments with minimum interruption to the system bal- 
ance. Backup relaying operates only in the event of a 
primary relaying failure; its action is only for uncleared 
faults. In mining, be it overload, short-circuit, or ground- 
fault relaying, both groups are used extensively to the 
point of redundance. 

Zones of Protection 

Protection to the entire system is principally related 
to primary relaying and is accomplished by establishing 
zones of protection. Each zone has an associated circuit 
breaker and fusible disconnect, or fuses with the required 
sensing devices, and adjacent zones overlap (10). If a 
failure occurs within an individual zone, only the switch- 
ing apparatus within that zone should open. If there were 
no overlap, there could be an unprotected region in the 
system in some situations; a failure within this area would 
produce no safety tripping. Although such a situation 
seems unlikely in theory, in practice it does occur and can 
be caused either by oversight or ignorance on the part of 
an engineer. It results in broad outages to the system, 
rather than restraining the problem within a specific zone. 



The coordination of protective relaying between the zones 
is extremely important. The aim is to isolate faults down- 
stream from the power source without disturbing up- 
stream zones. Unfortunately, obtaining this coordination 
is perhaps the most outstanding problem of relaying. 

By introducing the general arrangements of protec- 
tive relaying in resistance-grounded mine power systems, 
this section actually serves as a transition between the 
basic relaying principles and chapter 10, where there is a 
more detailed and specific analysis. The objective here is 
to show how primary and backup relaying, zones of 
protection, and coordination are utilized in both surface 
and underground mines. 

At all levels of the mine power system, protection 
against short circuits, overloads, and ground faults holds 
priority. Because the majority of failures in mining involve 
line-to-neutral faults, ground-fault protection commands 
special interest, but this does not negate the need to 
establish adequate line-conductor relaying. 

Coordination 

The objective of coordination is to determine the 
optimum characteristics, ratings, and settings for the 
protective-relaying devices (7); consequently, fault analysis 
of the system must be involved. The two common coordi- 
nation schemes that are utilized are pickup setting and 
time. With the first, relay pickup settings for a specific 
actuating quantity are set at progressively higher values 
from the loads to the power source, such that a higher level 
of the actuating quantity is required to trip the circuit 
breaker in an upstream zone. For the time coordination of 
a specific actuating quantity, pickup settings throughout 
the system will generally be the same, but the operating 
times to achieve contact closure at or above pickup are set 
progressively longer toward the source. One technique or 
the other may be applied to provide coordination between 
zones in the mine power system; at times, a combination 
might be needed. The design of this protective-relaying 
system can be a substantial problem, as exemplified by the 
fact that many engineers consider protective relaying 
more an art than a science. 

Ground-Fault Protection 

Except for capacitive ground current (see chapter 11), 
ground-fault current magnitudes are limited by grounding 
resistors. As mentioned in chapter 7, the maximum cur- 
rent is 25 A at low and medium voltages, but the level is 
seldom limited below 15 A, whereas it is very rare for a 
50-A limit to be exceeded on high-voltage grounding 
systems, with 25 A seen on many systems. Thus, ground- 
fault current is probably close to constant throughout the 
system. The use of delta-wye, delta-delta, or wye-delta 
transformers enhances this nearly constant current situ- 
ation. At each transformer step, a new or separately 
derived ground system is produced, each with its own 
ground resistor. Because the transformer configuration 
blocks zero-sequence components, a ground fault on cir- 
cuits connected to the secondary raises primary current, 
but the vectorial sum of the line currents is zero (1). 

It can thus be seen that coordination of ground-fault 
relaying across transformers is unnecessary, and both 
primary and backup protection must be established for 
each derived grounding system. Selective coordination at 
each voltage level by pickup setting alone is normally 



255 



impossible, and time settings must be relied upon when 
multistage protection is used (1). 

It must be remembered that resistance grounding is 
used to reduce fault energy, frame potentials, or system 
potentials. If the first ground fault is not cleared, an 
alternate ground fault in another machine will be in a 
ground condition and not be limited by the grounding 
resistor (1). The result is the possibility of dangerous frame 
potentials and powerful intermachine arcing. Conse- 
quently, ground-fault relays must always be arranged to 
trip circuit breakers; fuses cannot be used. 

Overloads and Short Circuits 

The occurrence of a line-to-line or three-phase fault 
anywhere in the system can cause substantial outages if it 
is not cleared downstream. Because they are not restricted 
like ground faults, the anomalous positive-sequence and 
negative-sequence currents can be passed across trans- 
formers to higher levels of the system. The resulting 
wide-range problems are especially evident on radial dis- 
tribution. System protection is usually coordinated by 
using instantaneous relays to adjust pickup, and the effort 
could involve fault currents from the machines to the 
substation. A maximum setting within an individual zone 
would be the minimum fault current at which thermal or 
mechanical damage to a protected device could occur. 

The problems resulting from inadequate overload 
protection are just as widespread. In most cases this 
protection must be applied to the specific zone that the 
fuse or circuit breaker is protecting. For example, the 
pickup setting of a time-delay relay could be determined 
by the ampacity rating of the smallest conductor within 
the zone of the switching apparatus, whereas time settings 
might be employed to coordinate between zones. Overload 
protection can be critical in large underground mines with 
numerous sections; when several machines are operating 
simultaneously, the maximum current demand could over- 
load an upstream conductor. 



KEY 

27 Undervoltage relay 
37 Ground-continuity relay 

50 ac instantaneous 

overcurrent relay 

51 ac time-delay 

overcurrent relay 

51G,51N ac time-delay ground 
overcurrent relay 

52 ac circuit breaker 
SA Surge arrester 



Frame, 
ground 




Grounding conductor 
Ground-check conductor 



52 

C U — (5 og) C } — fcoe) C D — (500) 



rr« 



£j- 



rr 

SJ6 



Excavator 



3f" 



T-©-! 



6 



Production 
shovel 



SWITCHHOUSE 



POWER CENTER 



Grounding^ 
resistor 






[ — 6oX 5 Sr — '.so'.si; 

)52 )52 

CD— |g) 



6 



Drill 



FT 

66 



Pumps 



Figure 9.60.— Simple surface mine power system illustrating 
protective relaying. 



Surface Mines 

To illustrate the application of these general consid- 
erations, consider figure 9.60, a one-line diagram (in- 
cluding grounding) of a simple surface mine power system 
(1). The diagram is drawn to show the various combina- 
tions of protective circuitry that can be found in surface 
operations. Each protective device is installed to trip the 
closest circuit breaker (52). 

In terms of ground-fault protection for the low-voltage 
machines, zero-sequence relaying (50G, usually instanta- 
neous) establishes primary protection. Backup protection 
could be placed here by adding timed direct relaying (5 IN) 
about the grounding conductor between the grounding 
resistor and the transformer neutral. For high voltage, 
remembering that the transformer configuration blocks 
zero-sequence current, primary ground-fault protection is 
provided in the switchhouse, again by zero-sequence relay- 
ing (50G, instantaneous or minimum-time dial setting), 
establishing a zone of protection for each outgoing circuit 
(the excavator, loader, and ac power center). The time- 
delay zero-sequence relay (51G) in the substation can be 
considered to give both backup protection for the down- 
stream 50G relaying and primary ground-fault protection 
for the zone between its location and the switchhouse. In 
any event, backup protection is the principal duty of the 



5 IN direct relay about the neutral conductor of the sub- 
station. (N is used to signify neutral relaying, but G can be 
used.) All this relaying is normally coordinated by time 
settings, the pickup setting being specified as a percent- 
age of the ground-current limit. 

The use of the grounding-conductor direct relaying, 
shown in figure 9.59, is restricted to situations where the 
main circuit breaker that the relay trips sees total ground 
current. In the ac power center, this means that the 
backup relaying must cause tripping of both circuit break- 
ers. Because here there is no main breaker, some States 
require potential relaying of the grounding resistor for 
backup. Potential relaying does give more safety than 
direct relaying, as previously mentioned. 

Ground-check monitoring is illustrated in its most 
extensive form for surface mining in figure 9.59, where 
every grounding conductor in the system is measured. 
Ground-check conductors are shown connected to each 
monitor. In instances where pilotless devices are applied, 
ground-check conductors are not needed. In cases where 
conductors from the substation form an overhead ring bus, 
such as in some open pit operations, ground-check moni- 
tors are often not used because of the circular nature of the 
distribution. In any overhead distribution arrangement, 
monitors can experience extensive failures because of 
lightning strokes. 



256 



Overload and short-circuit protection in the high- 
voltage portion is provided by line-conductor CT's usually 
connected in wye to time-delay relays that are normally 
induction-disk types with both instantaneous (device 50) and 
time-delay (51) elements. Each 50/51 combination estab- 
lishes a zone of protection downstream from its location. As 
shown in the power center, the same kinds of protection can 
be provided at utilization by low-voltage power circuit break- 
ers in conjunction with external CT's and relays. In figure 
9.60, these protection devices are indicated as numbers 
within dashed circles. Normally, they would be protected by 
molded-case units. Note that all 50/51 phase protection could 
also be done by fuses or fusible disconnects, but circuit 
breakers or power-driven load-break switches still are 
needed for ground-fault protection. 

Figure 9.61 shows a three-line diagram of a typical 
molded-case arrangement. These components replace the 
50, 51, 50G, and 27 devices associated with each low- 
voltage breaker in figure 9.60. The molded-case circuit 
breaker provides short-circuit protection through its mag- 
netic trip units (not shown), with overload tripping given 
by the internal thermal elements. (Overload protection 
may not be required on low- and medium-voltage circuits; 
see chapter 10 for discussion.) The external protective 
circuitry is commonly a zero-sequence relay and pilot-type 
ground-check monitor, and the contacts of both trip the 
undervoltage release. Notice that here the ground-fault 
relay shunts the trip coil, while the ground-trip relay is in 
series with the trip coil. The UVR itself provides under- 
voltage protection. In this way all the essential kinds of 
protection are available for a resistance-grounded trailing- 
cable installation. 

An undervoltage relay (device 27) is also shown in the 
switchhouse of figure 9.60, but this would not be required 
at this or any location provided that all equipment down- 
stream has undervoltage protection. 



Window 

(zero-sequence 

CT 




Receptacle -plug 
combination 



Figure 9.61.— Typical schematic for three-phase molded- 
case circuit breaker with ground-overcurrent and ground-check 
protection. 



Underground Mines 

For comparison, figure 9.62 illustrates a simple one- 
line diagram of an underground mine power system. The 
ac portion is very similar to the surface circuitry of figure 
9.60, and the reasoning and arguments behind the protec- 
tive devices are the same. There are differences, however, 
in the disconnect switch, the relaying in the substation, 
and possible extra outgoing circuits from the substation, 
but again, all these features are also possible in surface 
mines. 

By law, some means of visible disconnect is required 
within 500 ft of the point at which power enters the 
underground workings, and a separate switch is shown in 
the diagram for this purpose. Actually, it is also advisable, 
if not required, to place visible disconnect switches on 
incoming high-voltage (distribution) circuits within all 
power equipment. (These have not been included in figures 
9.60 and 9.62 merely to maintain clarity.) Consequently, it 
is not uncommon to find the first disconnect in the mine as 
part of a switchhouse. 

Ground-fault backup relaying located in this substa- 
tion is mainly by the potential technique. Here, the relay 
operating coil is placed directly across the resistor, and in 
addition to relaying, the transformer is used for resistor 
impedance matching for current limit. Although not 
shown, the outgoing circuit to the surface equipment has 
basically the same relaying as for underground (see chap- 
ter 13 for the implications of such loads). 

Direct Current 

Only one form of relaying for dc equipment is provided 
in figure 9.62; this is short circuit for rail haulage systems 
and consists of a dc overcurrent relay (device 76) driven by 
the voltage drop across an in-line shunt. For off-track dc 
machinery, the protective-relaying arrangement is di- 
rectly tied to the power source. Five basic systems of dc 
ground-fault protection are presently being used in the 
United States: 

• Diode grounding, 

• Basic grounding conductor, 

• Relayed grounding conductor, 

• Neutral shift, and 

• Differential current. 

The grounding philosophy for all but the last was intro- 
duced in chapter 7. The neutral-shift and differential- 
current systems can be used only when the machine is 
powered from the output of a rectifier, such as that 
contained in a section power center, whereas the first three 
systems are more commonly employed when a trolley 
system is the dc source. 

When dc equipment is powered from the trolley sys- 
tem, short-circuit and overload protection is normally 
provided by a dual-element fuse. The device is mounted in 
a holder or nip, which is clipped to the trolley wire. The 
other trailing-cable power conductor is connected to the 
rail. Fuses are rarely applied to trailing-cable protection 
in load-center rectifiers. Here, air-magnetic power break- 
ers, molded-case devices, or dc contactors (see chapter 12) 
are employed. In practice, these are usually tripped only 
for short-circuit conditions, with overload protection not 
used. 



257 




-ft-Chf--' h" 

] " r «j. £ © 

.'51k .»V / 



SURFACE AREA 

77777777777777777777777777777777777777777777777% 

UNDERGROUND 



KEY 

27 Undervoltage relay 

37 Ground-continuity monitor 

50 ac instantaneous 

overcurrent relay 

51 ac time-delay ground- 

overcurrent relay 
51G.5IN ac time-delay ground- 
overcurrent relay 

52 ac circuit breaker 
59G ac ground overvoltage relay 

72,72T dc circuit breaker 
76 dc overcurrent relay 



Machine frame 



MAIN 
SUBSTATION 



vf v Line-to-cable termination 
^7777777777777777777777777, 



SWITCHHOUSE 




MINER 



Figure 9.62.— One-line diagram of simple underground mine 
power system illustrating protective circuitry. 



Diode-Grounded 

The diode-grounded system is often found with dc 
vehicles that employ cable reels, because it permits a 
two-conductor cable to be used, which is less expensive and 
takes less space on the cable spool than type G cables. 
These features make the system attractive, but the diode- 
grounded system has deficiencies from a safety standpoint. 

A simplified diagram of the system is shown in figure 
9.63. The machine frame is tied to the grounded negative 
conductor by means of the grounding diode (DJ. The 
grounded conductor is connected to the power-center frame 
or trolley system rail, depending on the power source. In 
series with the diode lead is a ground-fault device, G, a 
mercury-magnetic switch or the operating coil of an elec- 
tromagnetic attraction relay. The pickup setting of this 
device can be no greater than 25% of the forward-current 
rating of the diode. 




Figure 9.63.— Diode-grounded system with possible fault in- 
dicated. 



If a positive-conductor-to-machine-frame fault occurs 
(location 2 of figure 9.63), current flows through the 
grounding diode and the ground-fault relay. When the 
fault current exceeds the pickup setting of the relay, the 
fault is then isolated by deenergizing the machine contac- 
tor (M x ) with opening of the G contacts. Actually, this 
sequence of events occurs only if the fault exists between 
the load side of the contactor and the motor. This leads to 
some of the basic drawbacks of the diode-grounded system 
since location 2 is the only safe area for a ground fault to 
occur. 

Location 1 of figure 9.63 includes the length of un- 
grounded conductor from its entrance point on the ma- 
chine to the main contacts of the M 1 contactor. Since these 
machines normally utilize cable reels, slip-ring assemblies 
are required for connecting the incoming power conductors 
to the machine circuit. Hence, it is difficult to locate a 
circuit-interrupting device at the immediate cable en- 
trance of the machine. If a fault occurs at location 1, the 
fault current flows through the machine frame, causing 
diode D x to become forward biased, and current then 
passes through the ground-fault relay. When the relay 
operates, the M x contactor resets, and the motor circuit is 
deenergized. However, this does not isolate the fault; 
isolation is solely dependent upon the opening of the M 1 
interrupting device, and the fault is located on the line 
side of its contacts. In this case, fault current can be 
terminated only by the switching apparatus on the source 
side of the trailing cable (dual-element fuses for a trolley 
nip, or fuses or a circuit breaker for a power center). 

Location 3 of figure 9.63 includes the entire length of 
the grounded conductor within the machine. A fault here 
renders the diode-grounded system unreliable, since it 
effectively provides a parallel path from the frame to the 
grounded conductor. The purpose of diode D 1 is to block the 
voltage drop across the grounded conductor from the 
machine frame. Therefore, if D x is bypassed, the frame 
potential is raised. A fault at location 3 can easily go 
undetected. As a result, a simultaneous fault at any 
location of the ungrounded circuit must again rely on 
isolation of the source side of the trailing cable. 

Failure of the grounding diode is a common problem 
associated with the diode-grounded system. With L\ in the 
shorted mode, the situation is the same as for a location 3 
fault. However, when the grounding diode fails in an open 
mode, all ground-fault protection is lost. Faults occurring 
in location 1 or 2 can cause the frame to become raised to 
a value equal to the supply voltage. 

The D 2 diode is referred to as the polarizing diode. Its 
purpose is to protect against installing the positive and 



258 



negative conductors in switched positions, which can occur 
when adding a new trailing cable or while making a cable 
splice. If the polarity of the supply voltage is inadvertently 
reversed, the polarizing diode will block the current flow, 
and the motor circuit cannot be energized. If D 2 failed in 
the shorted mode when the dc circuit was under reversed 
conditions, the diode-grounded system would be rendered 
totally inoperative, but the machine would still operate. 

Grounding Conductor 

Other methods of grounding dc machines require the 
use of a separate grounding conductor and type G cable. 
The basic grounding-conductor system (fig. 9.64) and the 
relayed grounding-conductor system (fig. 9.65) are the 
simplest of these techniques and are commonly used when 
the source is a trolley system. With the basic system, a 
ground fault is detected only by the overload protection on 
the source side of the trailing cable, and the result is 
extremely poor sensitivity with respect to ground faults. 
In the relayed system, the ground-fault sensitivity is 
improved by adding a ground-fault relay in series with the 
grounding conductor, as shown in figure 9.65. However, 
the current flowing through the relay may not be the total 
ground-fault current, since a parallel path may be estab- 
lished through the earth. 

Neutral-Shift System 

The neutral-shift system is illustrated in figure 9.66. 
The resistors Rj and R 2 create a dc neutral point for the 
system as well as limiting ground-fault current. A ground 
fault on either the positive or negative conductor will 
cause the neutral point to shift, which results in detection 
by the voltage-sensitive relays M 1 and M 2 . These relays 
are usually given the device number 64, a potential relay, 
also termed a dc unbalance relay. The primary limitation 
of this system is that it cannot discriminate between 
ground faults for individual pieces of equipment fed by the 
same rectifier. If a ground fault occurs on one machine, all 
circuits are interrupted. 

Differential Current 

As shown in figures 9.67 and 9.68, differential-current 
dc ground-fault protection utilizes the neutral of the 
wye-connected secondary feeding the rectifier. With a 
delta-connected secondary, a neutral connection can be 
provided through a grounding transformer. A grounding 
resistor is inserted in the neutral circuit to limit dc 
ground-fault current, typically to 15 A. Ground faults are 
then detected using either a saturable-reactor or 
saturable-transformer system. Actually, this results in an 
alternative method for grounding dc off-track equipment 
to that presented in chapter 7. 

With the saturable-reactor system (fig,. 9.67), a 
voltage-sensitive ac relay is placed in series with the 
winding of a reactor that encircles the positive and nega- 
tive outgoing conductors. The relay and reactor are excited 
by a 60-Hz control voltage, and under normal conditions, 
the current through the positive conductor equals that 
through the negative conductor. The magnetic fields about 
both conductors tend to cancel each other, and thus the 
reactor winding exhibits maximum impedance and pro- 
hibits the ground-fault relay from operating. However, 
when a ground fault occurs, the currents in the positive 
and negative conductors become unequal in magnitude 



Machine frame 




IX 

- v Grounding conductor 

Figure 9.64.— Basic grounding-conductor system. 






s 



Machine frame 



(+) \ // 




' Ground-fault relay 
Figure 9.65.— Relayed grounding-conductor system. 



To 
rectifier 



M< i 



-+• To circuit breaker 
-*• trip circuit 



(+) 



Mo) i 



<Po <-> 



GND 



To circuit breaker 
trip circuit 



Figure 9.66.— Neutral-shift system. 



because current is flowing in the grounding conductor. The 
magnetic fields about each power conductor are no longer 
equal, and the resultant magnetic field causes the reactor 
core to saturate. This in turn reduces the impedance of its 
winding, causing the ac control voltage to be impressed 
across the relay. Typical relay pickup is from 4 to 6 A of 
ground-fault current. 

The saturable-transformer system (fig. 9.68) operates 
on the same principle. Here, the ac control voltage excites 
the primary (bias) of a two-winding toroidal transformer. 
The transformer secondary is connected to a rectifier 
bridge whose output feeds an adjustable resistor in series 
with a dc relay. The relay has NO contacts when deener- 
gized. Under normal operation, ac is induced in the 
transformer secondary, and the rectified ac causes the 
relay to pick up its contacts. During a ground fault the 
core again saturates, but in this case the saturation stops 
transformer action. In turn, the dc is removed from the 
relay, which resets its contacts to cause circuit breaker 



259 



To circuit breaker 
trip circuit 



Grounding 
resistor 



To ac 
■►control circuit 




Figure 9.67.— Current-balance dc ground-fault relaying using 
saturable reactor. 



To circuit breaker 
trip circuit 




*■ Positive 

»■ Negative 

Ground 



To 
machine 



Figure 9.68.— Current-balance dc ground-fault relaying using 
saturable transformer. 



tripping. The resistor allows adjustment of the pickup 
setting. 

Another feature of differential current relaying in 
addition to sensitive detection of dc ground faults is that 
the technique senses dc unbalance only in the conductors 
that pass through the core. By using a detection device for 
each outgoing dc circuit, the individual relay system is 
sensitive only to ground faults existing downstream from 
its reactor or transformer. Thus, selective dc ground-fault 
relaying can be realized, an advantage not available with 
the neutral-shift system. 

The material in this chapter has served to introduce 
the complexities of protective equipment and relaying, 
culminating with figures 9.60 and 9.62, where it was 
shown that the protective relaying in a mine power system 
can be divided into zones of protection for the various 
power equipment. The application of this circuitry is the 
subject of chapters 12 and 13, where additional variations 
of these basic techniques are discussed. But first, a careful 
study of the overall mine power system is needed to 
emphasize the correct coordination of the various protec- 
tive devices. This is the topic of chapter 10. 



REFERENCES 

1. Chumakov, W. V. Protective Relaying for Mining Applica- 
tions. Pres. at 3d Annu. West Protective Relay Conf., Spokane, 
WA, Oct. 18-21, 1976; available from BBC, Allentown, PA. 

2. Cooley, W. L., and R. L. McConnell. Current State-of-the-Art 
in Ground-Check Monitoring. Paper in Conference Record -IAS 
11th Annual Meeting (Chicago, IL, Oct. 1976). IEEE, 1976. 

3. Cooley, W. L., and R. L. Rinehart. Grounding and Ground 
Check Monitors. Paper in Mine Power Distribution. Proceedings: 
Bureau of Mines Technology Transfer Seminar, Pittsburgh, Pa., 
March 19, 1975. BuMines IC 8694, 1975. 

4. Eaton, J. R. Electrical Power Transmission Systems. 
Prentice-Hall, 1972. 

5. Fink, D. G., and J. M. Carroll (eds.). Standard Handbook for 
Electrical Engineers. McGraw-Hill, 10th ed., 1968. 

6. Institute of Electrical and Electronics Engineers (New York). 
Recommended Practice for Electrical Power Distribution for In- 
dustrial Plants. Stand. 141-1986. 

7. Recommended Practice for Protection and Coordina- 
tion of Industrial and Commercial Power Systems. Stand. 
242-1986. 

8. Kaufman, R. H. The Magic of Pt. IEEE Trans. Ind. and Gen. 
Appl., v. 2, Sept./Oct. 1966. 

9. King, R. L. Development of an Electrical Engineering Course 
for Mining Engineers. M.S. Thesis, Univ. Pittsburgh, Pittsburgh, 
PA, 1977. 

10. Mason, C. R. The Art and Science of Protective Relaying. 
Wiley, 1956. 

11. Mason, R. H. Enforcement Begins for Ground Check 
Monitoring Underground. Coal. Min. & Process., v. 14, June 1977. 

12. Myers, W. P. Current-Limited Ground Fault Relaying. Min. 
Congr. J., v. 59, Apr. 1970. 

13. National Fire Protection Association (Quincy, MA). National 
Electrical Code. NFPA 70-1981 (ANSI Cl-1981). (Updated every 3 

yr-) 

14. Shimp, A. B., and D. A. Paice. Application of Molded-Case 
Breakers on DC Electrical Systems in Coal Mines. Paper in Mine 
Power Distribution. Proceedings: Bureau of Mines Technology 
Transfer Seminar, Pittsburgh, Pa., March 19, 1975. BuMines IC 
8694, 1975. 

15. Underwriters' Laboratories, Inc. Fuse Standards. UL 
198.1, 1973 et seq. 

16. U.S. Bureau of Mines. Schedule 2G, Electric Motor-Driven 
Mine Equipment and Accessories. Federal Register, v. 33, No. 54, 
Mar. 19, 1968. 

17. U.S. Code of Federal Regulations. Title 30-Mineral 
Resources; Chapter I -Mine Safety and Health Administration, 
Department of Labor; Subchapter O-Coal Mine Health Safety; 
Part 75 - Mandatory Safety Standards, Underground Coal Mines; 
Part 77 -Mandatory Safety Standards, Surface Coal Mines and 
Surface Work Areas of Underground Coal Mines; 1981. 

18. Wade, E. C. Ground Relaying for Mining Distribution 
Systems. Coal Age, v. 71, July 1966. 

19. Westinghouse Electric Corp., Low-Voltage Breaker Div. 
(Beaver, PA). Breaker Basics. 1973. 

20. Westinghouse Electric Corp., Relay-Instrument Div. 
(Newark, NJ). Applied Protective Relaying. Silent Sentinels Publ., 
1976. 

21. Wood, R. J., and H. D. Smith. Low-Profile Semiconductor 
Equipment for Mine Application. IEEE Trans. Ind. and Gen. Appl., 
v. 4, May/June 1968. 



260 



CHAPTER 10.— SIZING PROTECTIVE DEVICES 1 



In chapter 9, it was shown that circuit breakers, fuses, 
and switches are rated in terms of the nominal circuit 
voltage, the continuous currents they may carry, the 
sort-circuit currents they may interrupt, and the fault- 
through currents they must withstand, lb ensure that 
these interrupting devices disconnect faulted equipment 
promptly and correctly, it is necessary to have a separate 
protective system that recognizes the presence of a fault, 
determines what is faulted, and supplies energy to the 
mechanisms that will terminate current flow. It then 
becomes necessary to calculate the maximum fault cur- 
rents and, in many cases, the minimum sustained overcur- 
rent values in the mine system in order to determine the 
sensitivity requirements for the current-responsive protec- 
tion devices (3). 2 For multistage time-delay devices, the 
operating time of each device and the time relationships 
between devices must be found. These parameters are 
mandatory for the successful selection, installation, and 
coordination of protective equipment and relaying. 



FAULT CURRENT 

The fault current at any point in the mine complex is 
limited by the impedance of circuits and equipment from 
the source or sources to the fault point. The level is not 
directly related to the load on the system (5). When a mine 
is in development, system additions are often made to 
increase the capacity to handle the growing load. While 
these changes will usually not change the normal load on 
preexisting system portions, they may substantially in- 
crease short-circuit current during three-phase and line- 
to-line faults. Nevertheless, ground-fault currents, except 
for some special cases, remain relatively constant in 
high-resistance grounded applications. Whatever the situ- 
ation, the available fault currents must be predetermined 
to ensure adequate protective-device operation. 

Fault-Current Sources 

To find the available fault current correctly, all sources 
of fault current should be known. The main sources in 
mining are electrical utility systems, synchronous motors, 
induction machines, and capacitors (and capacitance). Syn- 
chronous generators are also a significant source, but these 
are only a concern if the mine generates its own power. Of 
the others, some can be eliminated because of their negligi- 
ble contribution to fault current, but with caution, depend- 
ing upon the installation. 

In most mine systems, the generators of the electric 
utility system are the principal source of short-circuit 
current. Because these generators are often remote from 
the mine, the current that results from a fault within the 
mining operation usually appears as merely a small 
increase in the generator load current, and the contribu- 
tion remains fairly constant (3-4). In this case, the electric 



1 The author wishes to thank Frederick C. Trutt, professor and chairman 
of electrical engineering, the University of Kentucky, for his assistance in 
preparing the fault-calculation and computer-analysis sections of this 
chapter. 

2 Italicized numbers in parentheses refer to items in the list of references 
at the end of this chapter. 



utility can be represented at the mine as a Thevenin's 
equivalent source— in other words, a single-value equiva- 
lent impedance driven by a constant voltage and referred 
to the point of connection. Even in proximate locations 
such as mine-mouth electric plants, this approximation 
usually gives adequate calculation results. 

Synchronous motors can be a substantial contributor 
to fault current, but because their prime coal mining 
application is as drive motors for excavators, they are 
mainly of concern to surface mines. The current supplied 
to a fault can be described as follows (3—4). After the 
occurrence of a fault, system voltage tends to decrease: the 
motor receives less power to rotate its load, and simulta- 
neously the inertia of the motor and load, as the "prime 
mover," causes the motor to act as a generator. The motor 
thus supplies fault current, which diminishes as the motor 
field excitation and kinetic energy decay. 

Induction motors can also contribute fault current to 
the system when inertia drives the machine as a generator 
(3-4). Here, however, the presence of field flux in the rotor 
is produced by induction from the stator; this decays 
rapidly with motor terminal voltage and disappears com- 
pletely after a few cycles. Accordingly, the effect of induc- 
tion machines, except for large horsepower ratings, can 
sometimes be neglected but should always be considered. 

Because of the widespread use of shielded cables, 
power-factor correction capacitors, and surge capacitors, 
capacitance is found in great abundance in mine power 
systems. Although the most destructive power-frequency 
fault currents come from rotating machinery (including 
the utility), capacitance can produce very high transitory 
overvoltages and fault currents (4). Fortunately, these are 
usually of short duration and of a natural frequency much 
higher than the power frequency, but they can lead to 
insulation damage. This subject is discussed in chapter 11. 

The stored charge of the capacitance acts as the prime 
mover, but because the fault-current contribution is of 
such short duration (less than 1 cycle), it can often be 
neglected when calculating line current (4). The main 
problem with fault conditions lies in the area of ground 
overcurrent during line-to-neutral faults. For instance, 
most high-voltage resistance-grounded distribution sys- 
tems have a 50-A ground-current limit or less, and some 
mines use a bank of surge capacitors across the terminals 
of all high-voltage loads. In 12.47-kV systems, the stan- 
dard 0.25-^F surge capacitor adds about 5-A capacitive 
ground-charging current, directly translating to ground- 
fault current during a line-to-neutral mishap. Adding to 
this current is the charging current of high-voltage feeder 
cables, typically 0.2 A per 1,000 ft. With just 10 capacitor 
banks on the distribution system in a moderately sized 
mine, a 50-A limit is exceeded. If a line-to-neutral fault 
occurs, the capacitance could discharge and feed the fault 
with capacitive current in excess of the ground-current 
limit. As this could happen within the distribution cir- 
cuitry, there is a possibility that the ground resistor would 
not see it, and because the duration could be less than 1 
cycle, the protective circuitry might not react. 

Source Equivalent Circuit 

The representation of equivalent circuits is the prin- 
cipal difficulty in the calculation of fault currents. An 



261 



equivalent circuit for the most important typical source, 
the utility system connection, has already been presented 
and is usually straightforward. Rotating machinery is less 
straightforward because the fault current contributed by 
each machine is limited by the machine impedance, which 
is unfortunately a rather complex and time-dependent 
variable. To simplify calculations of fault currents, 
Thevenin's equivalent sources are again assumed. 

Three specific values of reactance are used to estab- 
lish the fault current delivered at three points in time (3): 

1. The subtransient reactance, X^', for current during 
the first cycle after the fault occurrence; 

2. The transient reactance, X^, for current after sev- 
eral cycles at 60 Hz; and 

3. The synchronous reactance, X d , which determines 
current flow for the steady-state region. 

Substransient reactance is assumed to last about 0.1 s, 
after which the machine impedance increases to the 
transient reactance value. After 0.2 to 0.5 s, the value 
again increases to the synchronous reactance. With each 
increase, current contribution decreases. 

Depending on the maintenance of field excitation, 
synchronous motors could use all three values, but X d and 
sometimes X d are not needed in mining applications since 
the values approach infinity. Because of rapid current 
decay, X d is the only value used for induction motors, as 
these motors do not contribute significant fault current 
beyond the first cycle after the fault. Table 10.1 lists some 
typical motor reactances, given in per-unit based on the 
machine kilovoltampere rating (3). Following the previous 
statement and considering the type of rotating machinery, 
calculations on underground mining systems sometimes 
ignore the motor contribution, but computations for sur- 
face mines and surface facilities, especially where thermal 
dryers are involved, generally cannot. 



Table 10.1.— Sample reactances for synchronous and 
induction motors 





Subtransient 


Transient 




Motor type 


reactance 


reactance 


Comments 




(X"d) 


(X'J 




SYNCHRONOUS 








Motors: 








6 poles 


0.15 


0.23 ) 




8 to 14 poles .... 


.20 


.30 > 


None 


16 poles or more 


.28 


.40 J 




Converters: 








60O-V dc output 


.20 


01 
0/ 


None. 


250-V dc output 


.33 


INDUCTION 








Individual motors 


.17 


p) 


Motors generally above 
600 V. 






Groups of motors 


.25 


( 1 ) 


Motor voltage usually 600 


each less than 50 hp. 






V and below. 
Subtransient reactance 
is increased slightly 
because of very rapid 
current decay. Lower 
value of X" d would be in 
order for groups of 
larger motors. 



transient reactance not usually needed in calculations. 

NOTE. — The above values are in per-unit referenced to the motor kilovolt- 
ampere base. Approximate power bases can be determined from the motor 
horsepower rating: 0.8 pf motor, kVA base = hp rating; 1 .0 pf motor, kVA base 
= 0.8 (hp rating). 



FAULT CALCULATIONS FOR THREE-PHASE 
SYSTEMS 

In specifying equipment or checking protective- 
circuitry performance, certain simplifying assumptions 
are normally made to calculate fault currents. An impor- 
tant assumption when finding line currents on three- 
phase systems is that the fault is three-phase. This type of 
fault generally causes the maximum short-circuit current 
to flow. Perhaps the only exceptions are on utility systems 
and similar industrial applications where a line-to-neutral 
fault can cause current up to 125% of a three-phase fault 
(3), but these fault currents are substantially limited in 
most mining systems. The significance of assuming a 
three-phase fault is that symmetrical-component methods 
are not needed for the solution of routine fault calcula- 
tions, although detailed investigations may sometimes 
require looking at asymmetrical problems. 

Because mine power systems are normally radial or 
operated radially, calculating a three-phase fault is a 
rather simple task. Basically, all that is required is Ohm's 
law and an equivalent circuit. Even during a line-to-line 
fault, positive-sequence and negative-sequence imped- 
ances are essentially equal. A good estimation can thus be 
made by applying a fixed fraction of the three-phase case, 3 
which has been found to be about 0.87 (3). 

A second assumption is that the fault is customarily 
assumed to be bolted; that is, it has zero impedance (3-4). 
This not only simplifies calculations but also applies a 
safety factor, since the results provide a value greater than 
maximum and equipment is rarely stressed beyond its full 
rating. For instance, analytical studies on low-voltage 
systems have shown that minimum arcing fault currents, 
expressed as a factor of bolted faults, are typically 0.89 at 
480 V for three-phase arcing faults, and 0.74 at 480 V for 
line-to-line arcing faults (3). As the voltage level is in- 
creased, the arcing current level approaches but is always 
less than that of the bolted case. 

IEEE standards 141-1976 and 242-1975 detail exten- 
sive fault-calculation recommendations, many of which 
are directly adaptable to mine power systems (3-4). A 
summary of these follows, after which an example will be 
discussed. In both, specifics about sizing high-voltage 
switchgear will be presented to illustrate the method, but 
the same line of reasoning can be applied to low voltage 
and medium voltage. 

Short-Circuit Calculation Procedures 

The calculation procedure for finding currents result- 
ing from faults between power conductors, often termed 
short-circuit currents, can be divided into a series of steps. 
By assuming a bolted three-phase fault condition, the 
power-system parameters remain symmetrical regardless 
of the neutral conditions or even delta-wye transformer 
connections. The balance is especially close in high- 
resistance grounded systems, as line-to-neutral loads are 
not allowed. Therefore, the balanced fault currents can be 
calculated from a single-phase equivalent circuit, which 
has only per-phase impedance and line-to-neutral voltage. 
The procedure is to find the Thevenin's equivalent imped- 
ance of the entire system (Z), then solve for fault current 



3 Personal communication from E. K. Stanek, West Virginia University, 
Aug. 1977. 



262 



using Ohm's law. Calculations may use real, percent, or 
per-unit values. As shown in chapter 4, per-unit methods 
have the advantage of greatly simplifying the work when 
the system has different voltage levels. 

The first calculation step is always the preparation of 
a good one-line diagram that shows all fault current 
sources and all significant elements. All major imped- 
ances, both resistance and reactance, should be included: 
the utility, transformers, conductors, cables, and rotating 
machines. The collection of this information is perhaps the 
principal difficulty in fault calculations. Impedances for 
cables and conductors can be assembled from chapter 8. If 
transformer impedances are not known, the values can be 
estimated from the standard values given in tables 10.2 
and 10.3 (other impedances are given in chapters 12 and 
13) (3). 

Table 10.2.— Three-phase transformer per-unit impedances 1 
for liquid-immersed transformers, 501 to 30,000 kVA 

High-voltage Low-side rating, Low-side rating, 

rating, V 480 V 2,400 V and up 

2,400 to 22,900 0.0575 0.055 

26,400, 34,500 .0625 .060 

43,800 .0675 .065 

69,000 NAp .070 

115,000 NAp .075 

138,000 NAp .080 

NAp Not applicable. 

1 Actually a reactance, the value is normally expressed as a percentage. 
The per-unit values given use the transformer self-cooled kilovoltampere and 
voltage ratings as a base. Transformer resistance is usually well below 0.01 
pu. 



Table 10.3.— Three-phase transformer impedances for 
distribution transformers, including load centers 

High-voltage kVA Per-unit 

rating, V rating impedance 1 

2,400to 13,800 112.5- 225 >0.02 

300 - 500 >.045 

750 -2,500 .0575 

22,900 All .0575 

34,500 AH .0625 

Actually a reactance, the value is normally expressed as a percentage. 
The per-unit values given use the transformer self-cooled kilovoltampere and 
voltage ratings as a base. Transformer resistance is usually well below 0.01 
pu. 



In terms of system impedance, fault current has been 
shown to be mainly dependent upon the reactance be- 
tween the sources and the fault. This holds true except 
where there is substantial resistance such as with the 
extensive use of cables, overhead lines, and buses, which is 
the case in mining. However, if the reactance-to-resistance 
ratio, X/R, for the entire system from the source to the 
fault is greater than 5.0, negligible error is introduced by 
ignoring resistance. In fact, omitting resistance actually 
provides a small safety factor and has become common 
practice in other industrial applications. 

For most cases, it is recommended that the per-unit 
system be used in calculation. Therefore after the one-line 
diagram is complete, all parameters must be converted to 
per-unit values on a set of consistent bases. Normally, the 
base power, kVA b , is selected first. The current base, I b , 
and impedance base, Z b , are then derived, using the 
nominal voltage at each system level as the other base, V b . 



Accordingly, the base voltages must be related to the turns 
ratios of the interconnecting transformers. The power base 
may be any convenient level, but the main-substation 
kilovoltamperes is often selected. 

The next step in the calculation procedure is to reduce 
the system to its Thevenin equivalent by combining all 
impedances. The result of the reduction is a single driving 
voltage in series with a single impedance and the fault. 
Ohm's law can then be used to compute the per-unit fault 
current, or 



I -^ 

pu Z„ 



(10.1) 



where V pu = driving voltage of circuit, pu V, 

Z pu = equivalent impedance from sources to fault, 
including source impedances, pu fi, 
and I pu = symmetrical rms fault current, pu A. 

The prefault voltage, the level existing just prior to the 
fault occurrence, is assumed to be the nominal system 
voltage at the fault location. With this assumption, short- 
circuit current will approach maximum. Furthermore, 
when using the per-unit system, if the voltage bases are 
equal to the system nominal voltages, the driving voltage, 
V pu , is simply equal to 1.0 pu. 

The per-unit current can be converted to amperes using 
the base current multiplier. This calculated fault current is 
an alternating symmetrical quantity, because the sources 
are rms voltages, and can be used to compare equipment 
ratings that are expressed in symmetrical rms currents. 
However, the fault calculations must also recognize the 
asymmetry of typical fault currents and account for it. Fault 
current waveforms are discussed in chapters 4 and 9, but the 
typical asymmetrical type is again reproduced in figure 10.1 
for convenience (3). The compensation for asymmetry consid- 
ers current composed of two components: 

1. The ac symmetrical component, taken as the calcu- 
lated symmetrical value, and 

2. The dc component, with its initial maximum magni- 
tude taken as the peak of the ac symmetrical component. 



Total asymmetrical current 



dc component 



Symmetrical ac 




TIME 



Figure 10.1. 
metry. 



■Fault current waveform illustrating asym- 



263 



The time period in which the dc component decays is 
related to the reactance-to-resistance ratio, (X/R), of the 
system. Hence, the first cycle maximum fault current is 
estimated as 1.6 times the symmetrical rms value. De- 
pending on the X/R ratio, various other multiplying fac- 
tors are used to approximate maximum asymmetrical 
levels throughout the dc decay. These result in estimates of 
asymmetrical rms current that can be used to compare 
with ratings based on total (asymmetrical) rms current. 

Because fault current varies with time, several differ- 
ent results are commonly desired from the fault calcula- 
tions. To obtain these might require carrying out simulta- 
neous impedance reductions that account for X^', X^, and 
X d of the rotating machinery. However, when the system 
has just induction motors as loads, these usually contrib- 
ute fault current only within the first cycle, and at most 
only two reductions are necessary: one with the motors as 
sources, the other with just the utility system. In marginal 
situations, the motor contribution could be the extra 
current that results in equipment destruction during 
line-conductor faulting. 

Various applications of fault computations are shown 
in table 10.4. When aimed toward line-current applica- 
tions in mining, the first-cycle maximum symmetrical 
current is always needed. The maximum value that occurs 
between 1.5 and 4.0 cycles after the fault is used for sizing 
high-voltage circuit breakers. Fault current levels, possi- 
bly existing beyond 6.0 cycles, are needed to estimate the 
performance of time-delay relays and fuses. The calcula- 
tion result listed in table 10.4 refers to the ac symmetrical 
value of fault current. Figures 10.2 and 10.3, mentioned in 
the table, relate the system reactance-to-resistance ratio to 
the circuit-breaker contact parting time to obtain multi- 
plying factors for momentary or close-and-latch rating 
comparisons (3). To assist in using these curves, typical 
breaker speeds are 3, 5, and 8 cycles, with respective 
contact parting times of 2, 3, and 4 cycles (a 2-cycle circuit 
breaker has 1.5-cycle contact parting). 

As a summary of the foregoing procedures, the calcu- 
lation steps could be listed as follows: 

1. Express all impedances between the sources and 
the fault in per-unit on a set of consistent bases. (In rare 
cases where the sources and fault are at the same voltage 
level, such conversion is not necessary.) 

2. Reduce the system to one equivalent impedance. 



by 



3. Calculate the three-phase bolted-fault current, I s 



V 

T - P u T 

Z PU 



(10.2) 



where V pu = driving voltage of circuit, 1.0 pu, if nominal 
voltages are taken as the prefault level, pu 
V, 
Z pu = magnitude of equivalent impedance, pu fi, 
I b = base current for system portion in which 
fault exists, A, 
and I sc = symmetrical rms fault current, A. 

4. Apply appropriate multiplying factors to account 
for fault current asymmetry. 

5. If line-to-line fault current is desired, 



I P = 



I.-V3 



= 0.87 L 



(10.3) 



where I p = 



approximate symmetrical rms fault current 
resulting from bolted line-to-line fault, A. 



(Note that this is derived from the fact that positive- 
sequence and negative-sequence impedances are basically 
equal for phase-to-phase faults). For line-to-line-to-neutral 
faults, equation 10.3 also approximates the line-current 
level if the system is high-resistance grounded. 

Because of the high mobility of mine power equip- 
ment, any specific unit, say a switchhouse, could be 
installed anywhere, and the location is controlled more by 
the operation personnel than engineering personnel. It is 
often desirable that portable apparatus be able to with- 
stand or interrupt the worst fault conditions, and equip- 
ment should have a certain uniformity in design. To size 
this equipment correctly, repetitive fault calculations 
must be performed, assuming various fault locations 
throughout the mine system, to arrive at a worst case 
short-circuit current. More repetitions are usually needed 
at distribution levels than at utilization. Such extensive 
fault computations sometimes call for computer analysis, 
which will be discussed shortly. 



Table 10.4.— Sample applications of fault calculations 



Application 



Operation 



Fault current 



Comments 



1 . Fuses and low-voltage circuit 

breakers. 

2. High-voltage switching 

apparatus, excluding fuses. 



Contact parting and 
momentary withstand. 

Momentary rating or 
close-and-latch rating. 



3. High-voltage circuit breakers.... Interrupting duty. 



4. Time-delay relays Operation under 

short-circuit currents. 



First cycle maximum... Machine subtransient reactance, X" d , must be included. If ratings 
are symmetrical rms current, symmetrical fault current values are 
used directly. 

..do Machine subtransient reactance, X" d , must be included. Use 

values given in table 10.1, except use 1.2 X" d for individual 
induction motors 50 to 250 hp, neglecting those below 50 hp. 
Multiply symmetrical fault current values by 1 .6 to obtain total 
short-circuit current. 

From 1.5 to 4 cycles Multiply all motors X" from table 10.1 by 1.5, except use 3.0 X" d 
after fault. for individual induction motors 50 to 250 hp, neglecting those 

below 50 hp. Determine X/R ratio of Thevenin's equivalent 
impedance. Select appropriate multiplying factor from figure 10.2 
or 10.3. 

Beyond 6 cycles All motor contributions omitted. Procedure same as application 1. 



264 











3-phnsp 














140 




















x|<r 

o 120 
h- 






4 


3 2 


1 
















-TO-RESISTANCE R/> 

CD 00 O 

O o O 










' 






















/I 


s 










f 


>»l 












CD 1 
















REACTANCE 
o o 




















' 


7 




















' o^/ - 




















140 




























1 




120 


























/ 


I ] 


100 










~~1 


J1 












1 


7 < 


80 










J 


// 


















60 












/ t 

a 


>l 














40 




























20 








\rf>^ 










'co 


^ 


























1.4 1.6 1.0 1.2 

MULTIPLYING FACTOR 



1.4 



5 2 or more transformations 
(remote generation) 



4 Not more than 1 transformation 
(local generation) 

Figure 10.2.— Multiplying factors applied to three-phase 
faults to obtain momentary ratings for switching apparatus. 










-^n c 


o k 




f 


I 




I 


'I 




' 


C7 






o 




s 


7 




4j 












3-CYCLE 

- CIRCUIT - 

BREAKER 





1 


— t\j- 


f 




1 
a; 

>\ 

a> 
E 

Cn 

c 




1 / 

/ f 


1 








2-C 

- CIRC 

BRE 


rCLE 

:uit ■ 

iKER 



1.0 1.1 1.2 1.3 1.4 1.0 1.1 1.2 1.3 
MULTIPLYING FACTOR 
A Not more than I transformation ( local generation 



1.0 1.1 1.2 1.3 



1.01.1 




1.0 I.I 1.2 



1.2 1.3 1.4 

MULTIPLYING FACTOR 

B 2 or more transformations (remote generation) 

Figure 10.3.— Multiplying factors applied to three-phase 
faults to obtain closeand-latch ratings for switching ap- 
paratus. 



Three-Phase Calculation Example 

lb illustrate an example of bolted three-phase fault 
computations, consider the one-line diagram in figure 10.4, 
which approximates an underground coal mine in its early 
stages of development. All cables are given by their conduc- 
tor size; transformers by their voltages, capacity, and percent 
reactance; utility supply by its ability to deliver short-circuit 
current (a kilovoltampere capacity); and motors by their 
horsepower. Transformer resistance, power-equipment bus 
work impedance, and interrupting-device impedance are 
assumed to be negligible. Three possible fault locations are 
indicated, but for this example suppose only fault 1 has 
occurred. To ensure correct results, the contribution of rotat- 
ing machinery is included. 

The initial direction is to compute the first-cycle 
maximum fault current. Following the recommended pro- 
cedure, the first concern is to convert all impedances to 
per-unit. For calculation convenience, a base power, kVA b , 
of 5,000 kVA is chosen, a level not too large to make the 
per-unit values of any component insignificantly small. 
Prefault voltages are taken as nominal system voltages 
and define the second required base quantities, or for 
7,200-, 600-, and 480-V line to line converted to line to 
neutral, 

kV b69 = 39.838 kV, 
kV b72 = 4.16 kV, 
kV b 6 = 0.346 kV, 
kV b48 = 0.277 kV. 



Utility line 
1,000,000 kVA short-circuit level 



69 kV | A 3,000 kVA 

Z = 7.0% 

7.2kv" 



7.2 kvT M000 kVA 
ucuj Z = 5.0 /o 




5 g?°up P 100h P l00h P 50h P 



O O Q Q 




SURFACE 

. '= u ", h P FACILITIES 
induction 



SWITCH HOUSE 1 



600 
750 kVA v - l - u ^ 

7,200 I A 






1 2,000 
"3/C 4/0 
MPF 



SWITCHHOUSE 2 



-D — »- 



LOAD 
CENTER 



a 2,000' 
3/C 1/0 
MPF 



-a— * >■ 



; Fault 2 



Figure 10.4.— One-line diagram for fault calculations. 



265 



Calculations then provide the base current and base 
impedance for each base voltage. Thus, the base quantities 
for the 7,200-V system are 



At 600 V, 



At 480 V, 



kVA h 


= 5,000 kVA, 


kV h7? 


= 4.16 kV, 


^b7.2 = 


401 A, 


Zb7.2 = 


= 10.37 Q. 


kVA h 


= 5,000 kVA, 


kV b6 


= 0.346 kV, 


h.e = 


4,811 A, 


Zb.6 - 


0.072 Q. 


kVA h 


= 5,000 kVA, 


kV h48 


= 0.277 kV, 


Ib.48 = 


6,014 A, 


Zk 4.R = 


= 0.046 fi. 



The base impedance and current at 69 kV will not be used. 
The next process in the calculations is to reference the 
power-system parameters to these base quantities. The 
short-circuit level of 1 million kVA shown in figure 10.4 
relates to 1.0-pu impedance on the utility system at 69 kV. 
It is common practice to assume that this impedance is 
pure reactance, X pu , so the utility system base values can 
be taken as 



high side, and 69 kV, and is based on the transformer rated 
kilovoltamperes. The conversion is thus 

X % kVA b /kV e x 2 7(5,000) 

x - 2 = Took^ (kvj = Ir3o(3\oooj (1) = °- 117 pu - 



The 1,000-kVA load-center transformer has a percent 
reactance of 5% based on the rated capacity of 1,000 kVA 
and referred to the high side, 7.2 kV. Therefore, 



Xpu3 - 



5 /5,000 



100 \1,000 



(1) = 0.250 pu. 



The 750-kVA load center has a 4.5% reactance similarly 
referenced, so 

4.5 /5,000\ n nnn 

x p- = io1)(w) = - 300 p u - 



The resistance and reactance for each cable can be found 
from tables in chapter 8, and each impedance is easily 
changed to per-unit with 



Z -^ 

V - 7 



(10.5) 



kVA e = 1,000,000 kVA, 
kV e = 39,838 kV, 
Xpue = 1-0 pu. 

But the calculation base values at 69 kV are 

kVA b = 5,000 kVA, 
kV b69 = 39.838 kV. 

The utility per-unit reactance must therefore be converted 
using 



or 



_ X pue kVA b /kV e 
pu kVA b IkV 



(1.0X5,000) /39.8 



(10.4) 



^ = lLW00b-feJ =00 ° 5PU - 



Likewise, the per-unit reactance of transformers in the 
power equipment must be referenced to the calculation 
base quantities. The main substation contains a 3,000- 
kVA transformer, which has a 7% reactance referred to the 



Table 10.5 provides the results of all these cable computa- 
tions. 

The reactances in table 10.1 can be used to approxi- 
mate the variable impedance of each rotating machine. As 
a time saver, it is often best to determine all reactances at 
this point and carry them through the subsequent calcu- 
lations. However, because there are only induction motors 
in this example, subtransient reactance, Xj, is the only 
value defined, with transient and synchronous reactances 
approaching infinity. The load in the 480-V system is a 
150-hp group of induction motors with each motor less 
than 50 hp. Here the subtransient reactance is 0.25 pu, 
based on the motors combined with kilovoltamperes which 
is approximately the combined horsepower rating, or 150 
kVA. This reactance referenced to the base power for the 
calculations is thus 



xs = 



(0.25X5,000) 
150 



= 8.33 pu. 



The 600-V system contains a 500-hp motor group, two 
100-hp motors, and one 50-hp motor. Each motor in the 
500-hp group is about 100 hp. Hence, from table 10.1, this 



Table 10.5.— Impedance of cables in figure 10.4 



System, 
V 



Base 



Cable 



Actual impedance, ft 



Per-unit impedance 



impedance, fi Length, Size, 
ft AWG 



Type 



Voltage, 
kV 



Resistance (R) Reactance (X) Resistance (R) Reactance (X) 



7,200. 

480.... 
600.... 



Borehole. 



10.37 



.058 
.0752 



500 

2,000 

2,000 

1,000 

500 

500 

500 



4/0 
4/0 
1/0 
4/0 
4/0 
1/0 
2 



MP + GC 1 
MPF 
MPF 
G + GC 
G + GC 
G + GC 
G + GC 



0.032 
.126 
.254 
.068 
.035 
.064 
.109 



0.016 
.064 
.070 
.027 
.0135 
.0145 
.0145 



0.0031 
.0122 
.0245 

1.1724 
.4514 
.889 

1.514 



0.0015 
.0062 
.0068 
.4655 
.1875 
.201 
.201 



266 



subtransient reactance is about 0.17 pu on the combined 
kilovoltampere base, 500 kVA, and on the new base power 



Xa = 



(0.17X5,000) 
500 



= 1.7 pu. 



emphasizes the need to be cautious about neglecting the 
motor contribution. Nevertheless, this calculated value 
should be very close to the symmetrical component of fault 
current and can be used for time-delay relaying compari- 
sons (see application 4 of table 10.4). 



Continuing the process for the single-motor loads, the 
reactance for each 100-hp motor is 

X d ' = 0.17 pu on a 100-kVA base, 
X d ' = 8.5 pu on the 5,000-kVA base, 

and for the 50-hp motor, 



X d ' = 0.17 pu on a 50-kVA base, 

17 pu on the 5,000-kVA base. 



X d ' = 



Motor resistance could also be assinged, but since the X/R 
ratio for induction machines is about 6, this is neglected. 

The next step is to draw an impedance diagram as in 
figure 10.5 to show the calculated per-unit resistance and 
reactance values. The diagram should then be simplified 
as much as possible to show clearly the data required for 
calculations, as in figure 10.6. Even though this example 
concerns fault 1 only, all three fault locations are shown. 
All rotating machines, including motors and any genera- 
tors, are represented by their per-unit reactances, X d , X d , 
and X d , and are connected to an equivalent source bus (the 
dashed line in figure 10.6). Thus the utility supply equiv- 
alent reactance is in parallel with the rotating-machine 
reactances, and the combined potential driving the fault is 
1.0 pu. 

Further simplifications can now be made for each 
specific fault location. Figures 10.7 through 10.9 show the 
process graphically for fault 1. The operations involve no 
more than combining parallel and series impedances to 
obtain a single impedance between the source bus and the 
fault, which is the Thevenin's equivalent of the faulted 
system (fig. 10.9). 

The single equivalent impedance can now be used to 
calculate the first-cycle symmetrical rms current of the 
bolted three-phase fault, or 



■ V = 1.0pu 



0.005 pu 



. 0.117pu 



Only used to denote 
separate components 



: 0.0031 pu 



! 0.0015 pu 



: 1.478 pu 
; 0.587 pu 

■ 0.25pu 

X^' = 8.33pu 



\ 



fO.OI22pu 
i 0.0062 pu 



;; Fault! 



To signify that this is 
indeed a rotating machine 



0.0245 pu 
VWV 



0.0068 pu 



Fault 
2 




0.300 pu 




Fault 
3 



Figure 10.5.— Impedance diagram for one-line diagram of 
figure 10.4. 



I -S»I 



1.0 



0.1132 



(401) = 3,542 A. 



With the fault located in high-voltage distribution, this 
value can be used to compare fuse ratings based on 
symmetrical rms currents. If the fault had been in a 
low-voltage portion, the current would also apply to circuit 
breakers. It is obvious that in this example, resistance 
could have been omitted with negligible error. 

Figure 10.10A is a simplification of figure 10.5 to 
demonstrate the error that could be introduced by neglect- 
ing the motor fault current contribution. As shown, the 
only driving source is the utility, and the equivalent 
circuit in figure lO.lOfi can easily be found by series 
combination. The first-cycle symmetrical rms current in 
this situation is 






1.0 



0.1235 



(401) = 3,247 A, 



which is 295 A less than the previous value. Thus by 
omitting the motors, the result is an 8.3% error, which 



Shows source of 
driving potential 




X Fault 
2 



0.0245 pu 0.3068 pu 

Figure 10.6.— Simplification of figure 10.5. 



267 



If fault calculations are to be applied to high-voltage 
switchgear, the momentary or close-and-latch current du- 
ties must be calculated. The next step in the calculations 
would therefore be to account for fault current asymmetry 
by employing multiplying factors that adjust for the dc 
component decay. Here the subtransient reactance of all 
induction motors greater than 50 hp is modified by mul- 
tiplying by 1.2. The contribution of motors less than 50 hp 
is neglected. These adjustments can be compared to a 
decrement factor, allowing for fault current decay with 
time. Figure 10.1 LA reflects these changes and also omits 
resistance (compare with figure 10.6). Figure 10.1 IB 
shows the reduced single equivalent reactance, and using 
the formula, 



V pu 

Isc(mom) — rj lb d-6) 



(10.6) 



0.122 



0.0031 



0.0015 



1.478 



. 9.167 X d 



; 0.0367 ,( 



0.313 



Fault 



-Note : fault 2 removed 



• 0.486 > 0.444 

1.888 X d " C 4.35 X d " C 17.201 X d 



Figure 10.7.— Simplification of figure 10.6. 



or 



Isc(mom) = OHM (401X1.6) = 5,559 A. 



This current represents the maximum asymmetrical rms 
value in the first cycle and can be compared with close- 
and-latch capabilities of high-voltage switching appara- 
tus. These are normally circuit breakers, but any device 
that needs to be rated on total fault current also applies. 
Interestingly, if subtransient reactances were not modified 



V = 1.0pu 




Figure 10.9.— Equivalent circuit of figure 10.6. 



V=1.0pu 



J0.I22 

j 0.0031 
. j 0.0015 

Fault 



A Utility, substation, and 
borehole cables 



V = 1.0pu 




Figure 10.10.— Example problem with motor contribution 
neglected. 




■=> 



■ 0.0033 



0.1219 



: 0.2824 X Fault 



1.543 



Figure 10.8.— Further reduction of example network. 




A Complete network 



V-I.Opu 



J0.II54 



B Equivalent circuit 



Figure 10.11.— Network to calculate momentary or close- 
and-latch current duties. 



268 



in this example, the calculations would produce an asym- 
metrical rms current of 622 A, thus creating a small safety 
factor for switching-apparatus comparison. 

To illustrate a further application of fault calculations 
(table 10.4, item 2), the circuit of figure 10.6 can again be 
adjusted to determine the interrupting duties of high- 
voltage circuit breakers that have minimum contact part- 
ing times between 1.5 and 4.0 cycles. Again, to allow for 
decrement of the dc component, subtransient reactances 
are multiplied by (3) 

1.5, for synchronous motors, 

1.5, for 3,600-r/min induction motors over 250 hp and 
1,800-r/min induction motors over 1,000 hp, and 

3.0 for induction machines of 50 hp and larger (item 2 
supersedes this item where there is any overlap in coverage). 

Smaller motors are considered insignificant. To obtain the 
necessary multiplying factor from figures 10.2 or 10.3, the 
X/R ratio of the equivalent impedance or fault-point im- 
pedance is needed. This can be found by reducing 
resistance-only and reactance-only networks, which elim- 
inates handling complex numbers. The formula for inter- 
rupting duty is then 



where 



Iscont) = if* h (multiplying factor), 



L 



(10.7) 



j. sc( tl = total rms current interrupting 
duty, A, 
and multiplying factor = value from figure 10.2 or 10.3. 

For comparison with circuit breakers rated in megavolt- 
amperes, the interrupting capacity can be converted to an 
interrupting-current capability by 



(10.8) 



MVA 
Isdint) - ^3 y , 

v <J Y oper 

where I sc(int) = interrupting current, kA, 

MVA = system power, MVA, 
and V r = line-to-line system voltage, kV 



The system voltage is used when it is within the minimum 
and maximum rated limits of the breaker, but the minimum 
rated breaker voltage is used for formula 10.7 if the operat- 
ing voltage is less. For instance, if the minimum rated 
voltage is 12.47 kV and the circuit breaker is being used on 
a 7.2-kV system, then 7.2 kV is used for calculations. 

The foregoing example is just a simple demonstration 
of bolted three-phase fault calculations in a small mine. As 
a result, the fault current was rather small but still large 
in comparison with normal load current. In practice, more 
fault locations would be taken, and the computations 
repeated until the worst case condition was verified. As 
more mining sections or loads are involved, the complexity 
of these calculations increases substantially, oftentimes to 
the point where hand computing becomes too cumbersome 
and a computer analysis must be adopted. 

Computer Fault Analysis 

Systematic digital computer methods, developed to 
aid the designer in specifying system protection require- 
ments, are usually called upon for major fault studies (3). 



A complete discussion of the theoretical bases of such 
techniques is beyond the scope of this chapter, but is 
available in the literature (for instance, reference 15). 
However, a few thoughts about the methods follow. 

Balanced-fault calculations using a computer require 
either a generalized impedance formulation of the power- 
system network (referred to as ZBUS fault analysis where 
the network is similar to the preceding calculations) or a 
generalized admittance formulation, which is termed 
YBUS load-flow analysis. The impedance network has the 
advantage of also allowing the evaluation of 
unsymmetrical-fault behavior by applying symmetrical- 
component theory. With this approach, line-to-neutral, 
line-to-line, and line-to-line-to-neutral faults may be sim- 
ulated by proper connection of the power-system network 
representation for the positive, negative, and zero se- 
quences. On a large digital computer, computations are 
quite rapid, and the amount of output information is 
limited primarily by the user's requirements. Available 
information generally includes fault currents, as well as 
postfault currents and voltages, for any fault at any 
system location. Information on sequential faulting of any 
system bus or location is possible. Prefault load currents 
are normally neglected but may be included if a more 
exact analysis is required. 

Input-output techniques vary with the program type 
but may be either batch or interactive mode, and in either 
case the system is generally input in terms of impedance 
or even cable sizes between bus pairs. Prefault loading 
information may also be input in terms of kilovoltamperes 
or load type. In batch-processed programs, the run is then 
completed and required output information displayed. 
Interactive techniques allow the additional capability of 
on-line system modifications to evaluate the effects of 
design changes (7-8). 

Ground-Fault Current Calculations 

Federal regulations require high-resistance ground- 
ing on portable or mobile coal mining equipment. For 
low-voltage and medium-voltage systems, the neutral 
grounding resistor must limit the maximum ground-fault 
current to no more than 25 A. For high-voltage systems, 
the voltage drop in the grounding circuit external from the 
resistor must be 100 V or less under grounding-fault 
conditions. To meet these conditions, as well as to exceed 
capacitive charging-current constraints, high-voltage min- 
ing systems typically have maximum ground current 
limits at about 25 A, but rarely above 50 A. As the 
grounding conductor cannot carry load current, there can 
be no neutral-connected loads in these systems; thus, 
ground-fault current is contributed only from generating 
stations and capacitance. In other words, motors have no 
effect. Power-factor correction capacitors are not connected 
to the neutral, so capacitance is contributed only from 
conductor and cable capacitance and surge capacitors. 
This contribution can be substantial, but it is always less 
than the maximum ground current limit and is of very 
short duration, much less than 1 cycle. 

Under these conditions, line-to-line-to-neutral faults 
cause about the same current level as line-to-neutral 
faults, and the symmetrical rms ground-fault current can 
be calculated from the line-to-neutral case, or 



! = 3^ ^ 



(10.9) 



269 



where l^ = symmetrical rms ground-fault current, A, 
V Ln = line-to-neutral voltage of system, V, 
Z x = positive-sequence impedance, 12, 
Z 2 = negative-sequence impedance, Q, 
Z = zero-sequence impedance, Q, 
Z G = sum of impedances for fault and grounding 
circuit, fl, 
and R G = resistance of neutral-grounding resistor, Q. 

The sequence impedances are usually so close that they 
can be assumed equal, so that 



Ief = 



3V T 



V, 



3Z T + 3Z G + 3R G 



Zjrp + Lirz + Iw 



(10.10) 



The sequence impedances (Z T ) are usually so small that 
they can be neglected, and 



T n 

" ~ Z„ + R r 



(10.11) 



Consequently, the calculation of ground-fault currents on 
high-resistance-grounded systems is a rather simple task. 
To obtain maximum values, many engineers assume a 
bolted fault and sometimes neglect conductor impedance. 
For minimum available ground-fault currents, conductor 
impedance can be the maximum possible at the mine, but 
a bolted fault is again usually assumed. 

This simplification does not apply to grounding sys- 
tems that are not high resistance grounded, but equation 
10.9 can still be used to approximate ground currents from 
line-to-neutral faults. The method of symmetrical compo- 
nents should always be used in these asymmetrical fault 
cases. 

DIRECT CURRENT SYSTEM FAULTS 

The preceding material covered ac occurrences almost 
exclusively, yet knowledge about fault current levels on 
the dc portions of mine power systems is also imperative 
for the correct application of protective equipment, since 
faults can occur within motors, on trolley lines, trailing 
cables, cable reels, surface-excavator loop circuits, and so 
on. Trolley lines can pose extensive problems because of 
their normally heavy load currents as well as their ex- 
posed nature; for example, a roof fall can cause a serious 
arcing fault. 

In underground mines, the principal source of dc fault 
current is from the dc distribution, through rectifiers and 
to a lesser extent batteries (see chapter 15) and motor- 
generator (m-g) sets. For surface mining, Ward-Leonard 
systems are also involved, but these fault locations are 
commonly confined to inside unit-designed equipment. As 
with ac, the voltage drop associated with a fault can cause 
motors to go into a generation mode and act as a source as 
long as the field and the applied inertia are maintained. 
Prime movers are perhaps a more serious concern with dc 
than with ac since they can exist for an extended time, for 
instance, the inertia of a large trolley locomotive being 
pushed by loaded cars. However, the motor contribution 
varies widely from application to application, so that 
assigning values for series impedances, like subtransient 
reactances, can be a very difficult task. Furthermore, 
procedures for simplifying calculations are not well estab- 
lished, and many authors recommend very detailed tech- 
niques (1, 3). 



Fortunately, dc systems in mining are almost exclu- 
sively at utilization, and in them the fault current seen by 
protective devices is usually dependent upon location and 
not motor contribution. In other words, a motor might 
pump substantial current to the fault but not usually 
through a protective device unless the fault is upstream 
from the device. Thus, in many cases, the fault network 
can be effectively constructed using only the principal 
sources and the impedance in series with the fault. When 
basing the reasoning on this apparent simplicity, it is all 
too easy to treat dc faulting as just an application of Ohm's 
law and resistance and fail to recognize that the system 
current is no longer steady state after the fault. Figure 
10.12 shows an example in which the fault current is 
interrupted before the maximum is reached (14). It is 
apparent that the rate of rise as well as time to interrup- 
tion must be considered, and circuit inductance becomes a 
very important factor (1). In assembling a fault network, 
trolley-system inductance is not an easy parameter to find, 
but values between 0.2 and 2.0 mH have been found to be 
typical (14). 

Figure 10.13 gives plots for some rectifiers versus 
distance of the fault from the rectifier (14), to illustrate the 
available fault current level that can exist on rectifier-fed 
trolley systems. One of the more prevalent dc protective- 
relaying problems in mining is not maximum fault level 



or 
or 

o 
en 

LU 

< 
U 
C£ 
CD 






prospective 



Fault 

occures 




TIME 



Figure 10.12.— Fault current in dc system. 





40 


iii i | i i i i | i i i 
^ KEY 


i 


< 




\ 1,000-kW, 300-V rectifier 




ro 
O 




1 1,000 -kW, 600- V rectifier 




\-~ 


30 


-\ 500- kW, 300-V generator 


— 


UJ 

rr 

O 


20 


I _. — 500-kW, 300-V generator 
, 1 1 




Ll) 
> 

(- 

o 

UJ 
CL 
(/) 
O 

rr 


10 




- 


LL 


i i i i 1 i i i i 1 i i i i 



5 10 

DISTANCE FROM SUBSTATION. 10 3 A 



15 



Figure 10.13.— Available fault current versus distance of 
fault from rectifier on typical trolley systems. 



270 



but rather the possibility of having small undetected 
faults occur on trolley systems. In Pennsylvania, during 
1978 alone, two extensive mine fires were initiated in this 
way. A trolley-wire-to-rail fault, for example, could occur as 
the result of a roof fall but be of high enough resistance so 
the fault current would be less than normal load currents; 
hence, the fault would not be detected by conventional 
overcurrent relaying. Several methods have been proposed 
to detect such illegitimate loads, including discriminating 
relaying and rate-of-current-rise detection (11). At this 
writing, these are still in the demonstration stage. 

DEVICE SETTINGS 

Coordination of a mine power system entails complete 
organization of time settings and/or current settings for 
all protective devices from the loads to the sources. This 
necessitates a comprehensive coordination study of the 
entire system to determine the range of correct values for 
all instrument transformers, pickup and time settings, 
fuse ratings, and circuit breaker trip ratings, which will 
provide effective coordination and selectivity and ensure 
that the minimum of unfaulted load is disturbed when 
protective devices isolate a fault. The prime concern is 
overcurrent, since the circuitry must provide simulta- 
neous overload, short-circuit and ground-fault protection 
without causing nuisance tripping. The application of this 
"art" is perhaps the most perplexing problem facing 
practicing engineers. A system fault analysis is vital input 
for any comprehensive coordination study and should 
address not only maximum values but also minimum 
values, together with the normal operation and maximum 
allowable currents. 

The balance of this chapter is broken into the major 
aspects of the coordination study: relay pickup settings, 
CT matching, circuit breaker trip settings, fuse character- 
istics, and overall coordination. Emphasis is placed on 
radial ac systems that are high-resistance grounded. 

RELAY PICKUP SETTINGS 

Pickup has already been defined as the minimum 
value of the actuating quantity that will cause a relay to 
operate its contacts. Whether the application is at a 
trailing cable, feeder cable, or overhead conductor, the 
requirements for overload, short-circuit, or ground-fault 
protection usually translate into current pickup settings. 
The values used in this chapter are generally in line with 
those contained in 30 CFR 75 and 77, which are in effect 
at this writing (16). The quantities that define pickup may 
change in the future, but the techniques presented here 
for establishing relay pickup settings should not. Because 
of their widespread use, induction-disk relays are implied 
in most of the following pickup applications; however, 
pickup techniques for other relays are basically the same. 
To avoid confusion, molded-case circuit breaker trip set- 
tings will be covered later. 

Establishing a pickup setting for an ac relay involves 
selecting a CT ratio and operating-coil current. For mar- 
ginal short-circuit currents, the accuracy of the combina- 
tion might require verification. The following material 
describes pickup settings for a single zone in a mine power 
system, but it must be remembered that the overall goal is 
to obtain coordination, and the settings at any location can 
be affected not only by requirements and regulations but 
also by other upstream and downstream relaying. 



Short-Circuit Protection 

Short-circuit protection can be obtained with an in- 
stantenous element (no intentional delay) or an inverse- 
time overcurrent relay using the minimum time dial 
setting (maximum time delay here is often restricted to no 
more than 0.6 s). The general requirements can be deter- 
mined by selecting the lower value calculated from the 
following: 

1. 115% of the maximum starting current or 115% of 
the peak load current, whichever is higher, for the equip- 
ment being protected; or 

2. 60% of the smallest bolted three-phase symmetrical 
rms fault current for any point of the zone protected by the 
relay. 

The first value is designed to have pickup above the 
normal operating current to prevent nuisance tripping. 
Usually, the bolted fault current value is higher than the 
first value. 

If the maximum starting currents of the motors are 
not known, each can be approximated by 



L = 1.25 



x;') If 



(10.12) 



where I s = approximate maximum starting current, A, 

= per-unit subtransient reactance of motor or 
motor group, pu Q, 
and I n = motor full-load current, A. 



X3 



The motor full-load current can be estimated from 



746(hp) 
lfl " V3 Vi; (pf) ' 



(10.13) 



where hp = rated machine horsepower, 

V = rated line-to-line voltage of motor, V, 
rj = motor efficiency, 

and pf = full-load power factor, which can be assumed 
to be 0.85. 

Inrush currents for transformers usually range from 8 to 
12 times the full-load current rating for a duration of 0.1 s; 
typical values for inrush currents should be available from 
the transformer manufacturer. 

To show how short-circuit pickup is selected, consider 
that a production shovel in a surface mine has 2,000-hp 
connected load rated at 4,160 V line to line. As shown in 
figure 10.14, the shovel is powered through 1,000 ft of 4/0 
AWG cable. The shovel induction motor is 85% efficient at 
full load and operates at 0.8 pf. The minimum value of 
bolted three-phase fault current has been found to be 6,130 
A. CT ratios are 400:5 A, and the instantaneous element 
has a pickup range from 10 to 50 A. The procedure is to 
find the minimum current according to the above criteria. 

First, for the fault current, the pickup would be 60% of 
the bolted value divided by the CT turns ratio: 

(0.6X6,130) AC . 
pickup = oq = 4b A. 



This value must now be compared with the pickup needed 
to slightly exceed maximum starting current, which may 



271 



be estimated using equations 10.13 and 10.12. Here, the 
full-load current of the shovel motor is about 



_ 746(2,000) 

^ ~ V3 (4,160X0.85X0.8) 



= 305 A. 



From table 10.1, the per-unit subtransient reactance of 
individual induction motors is 0.17, and the estimated 
maximum starting current is then 



I s = 1.25 [——J (305) = 2,243 A. 



50/51 

o 



> — (f 

4,160-V, 
6,130-A 
symmetrical 
rms fault 
current 



52 

-D- 



y^K( 



I.OOOft 4/0 SHD-GC 



«r 



o 



2,000-hp 
shovel 



Figure 10.14.— One-line diagram for pickup setting example. 



Allowing 115% to prevent nuisance tripping, the pickup 
setting is then 



• i (2,243X1-15) QQ 
pickup = on = 32 A. 



As this is less than the value for the fault current, 32 A is 
the selected pickup setting for short-circuit protection. 

Overload Protection 

The prime purpose of overload pickup settings is to 
protect conductors and insulation from damage by excess 
temperature. Temperature here is a function of the ambi- 
ent temperature and the I 2 R power loss in the conductors; 
temperature rise also involves time; therefore, inverse- 
time overcurrent relays are used on high-voltage systems 
to provide this protection. (The thermal overloads and 
fuses commonly used for low-voltage and medium-voltage 
systems will be discussed later.) The general overload 
requirement is that relay pickup should occur whenever 
125% of the ampacity of the smallest power conductor in 
the protected zone is exceeded. 

When this requirement is directly applied to compo- 
nents such as transformers, the overload value could be 
125% of the rated full-load current. However, the percent- 
age recommended for transformer overload protection var- 
ies depending upon the protection scheme associated with 
the transformer. For transformers rated greater than 600 
V, the National Electrical Code allows 300% of rated 
primary current with circuit breaker protection, and 150% 
for fuses when there is no protection at the transformer 
secondary (10). With transformers rated less than 600 V, 
overload protection for the primary winding is 125% when 
there is no overload protection at the secondary and 250% 
if overload protection for the secondary is set at 125% (8). 
On the other hand, IEEE standard recommendations for 
transformer overload protection state that time-relay 
pickup should be set at 150% to 200% of the primary 
full-load current (3). 

Considering figure 10.14 as an example for overload 
pickup, the trailing cable is 4/0 AWG, three-conductor, 
5-kV, with 90°C rated insulation, and the relay operating 
coil (for the 51 contacts) has a pickup range of 4.0 to 12 A. 
The maximum ambient temperature at this location is 
30°C. From the data given in chapter 8, the cable ampac- 
ity at 40 °C is 321 A, and when applying the temperature 
correction factor for 30°C, 

ampacity = (1.1X321) = 353 A. 



Allowing for 125%, the minimum conductor current that 
defines overload is 

I = 1.25(353) = 442 A. 



The CT ampere-turns is 400:5 A, so the required pickup 
setting is 

442 
pickup = -qtt = 5.5 A. 



A tap setting corresponding to this pickup would allow the 
time-current characteristics of the relay to determine an 
overload condition. If the next tap setting available was 
6.0 A, this would relate to 480-A conductor current as an 
overload, which is too high. If an instantaneous element 
with a 32-A pickup is available in the relay, short-circuit 
protection is also provided with one CT per line conductor. 

Ground-Fault Protection 

Most experts suggest that ground-fault relays should 
pick up at no more than 30% of the maximum current 
limit of the grounding system to provide an ample margin 
of safety in high-resistance grounded systems (9, 13, 19). 
For a 25-A current limit, this represents a line-conductor 
unbalance producing a zero-sequence current of about 8 A. 
However, such a demand is lower than present protective- 
circuitry detection capabilities when electromechanical 
relays are used. For instance, the optimum arrangement 
with induction-disk relays is zero-sequence circuitry with 
a 25:5 ampere-turns CT in which the most reliable pickup 
performance is not less than 12 A. One reason for this 
limit is connected to the fact that the window-type CT 
needs a large opening in order to pass the three line 
conductors through. Here, zero-sequence currents less 
than 12 A do not generate enough capacity from the CT to 
drive the relay adequately. 

Another problem is that induction-disk relay opera- 
tion is not reliable when the magnitude of actuating 
current is only slightly above the tap setting. This is 
because the net actuating force is so low that any addi- 
tional friction in the rotating-disk mechanism can prevent 
operation or increase the operating time. Even if the relay 
does close its contacts, the contact pressure may be so low 
that contamination of the contact surface can prevent 
electrical contact. To minimize this problem, it is common 
practice to apply induction relays in such a way that their 
actuating quantities are at least 1.5 times the tap setting 
(6). In fact, time-current curves are rarely shown for less 
than this amount. On the other hand, an induction relay is 



272 



most effective if its pickup is selected so it will operate on 
the most inverse part of its time curve. Thus, the mini- 
mum value of actuating current should be only slightly 
higher than 1.5 times the tap setting. Often a pickup 
corresponding to 2.5 times is selected for very inverse 
relays, the type usually applied to ground-fault protection. 

It might be thought that requiring relay operation 
below the ground current limit is too stringent and that 
relays should only be required to operate at the fault 
current available on the system. This is a logical deduc- 
tion, which relates that on a properly installed grounding 
system no potential greater than that allowed can ever 
exist between any metallic object and earth under fault 
conditions (40 V on low-voltage and medium-voltage sys- 
tems and 100 V on high voltage). Nevertheless, in terms of 
personnel safety, setting a relay pickup at less than the 
limited fault current provides viable backup protection. In 
other words, machine frames would remain well below the 
current allowed during any ground fault. Any trend to- 
ward hazardous conditions could also be detected, al- 
though such trends involving the system neutral are 
infrequent on high-voltage distribution systems. 

Considering these thoughts and in-mine practice, 
ground current pickup should not be greater than 50% of 
the current rating of the grounding resistor. This level 
should provide reliable repetitive operation of zero- 
sequence protective circuitry. As technology improves, 
30% pickup should be the goal, as is suggested in the 
literature (9, 13, 19). 

To give an example of pickup settings, consider that 
ground-fault protection is provided by zero-sequence relay- 
ing A very inverse induction-disk relay is used and has a 
tap-setting range of 0.5 to 4.0 A for the time element coil. 
The CT has a 25:5 ampere-turns ratio, and the ground 
current limit is set at 25 A. Applying the 50% recommen- 
dation, zero-sequence current flow in the three line con- 
ductors cannot be greater than 12.5 A for relay pickup; 
thus, the maximum tap setting of the operating coil is 
12.5/5 or 2.5 A. However, the selected tap setting should be 
lower to allow for effective time-delay operation. Using 2.5 
times the minimum value of actuating current, the tap 
setting for the relay should be 2.5/2.5 or 1.0 A. 

On solidly or low-resistance grounded systems, the 
available ground-fault current is substantial enough that 
there are usually no relay pickup problems, even if resid- 
ual ground-fault relaying is used. Accordingly, the pickup 
level in this case could be defined by 30% of the minimum 
bolted line-to-neutral fault current. Similar logic could 
also apply to ungrounded systems, but here the maximum 
ground current requirement or even a ground-fault re- 
quirement is not easy to define. As a result, ground-fault 
pickup could, if necessary, be related to a decrease in the 
line-to-neutral potential of any power conductor and per- 
haps be 30% of the nominal system voltage. This line of 
reasoning could apply to dc as well as ac systems. 



CURRENT TRANSFORMER MATCHING 

From the foregoing, it appears that the selection of CT 
ratios and relay pickups to provide protection against 
excessive currents is quite a straightforward process. How- 
ever, CT's are magnetic devices and can give inaccurate 
results when improperly applied. CT performance is an 
important factor in protective-relay design because relays 
are only as accurate as the CT's that energize them. The 
main CT problem is core saturation caused by excessive 



primary current, incorrect secondary burden, or a combi- 
nation of both these factors {4). When a CT is in satura- 
tion, its accuracy is very poor, secondary current is actu- 
ally less than it should be, and relays tend to operate more 
slowly than intended. One danger is the loss of relay 
coordination. 

Current Transformer Accuracy 

A model of a CT and its burden is shown in figure 
10.15. The dependent current source delivers a current 
equal to the primary current (I p ) divided by the turns ratio 
(N) of the CT. The remaining impedances, voltages, and 
currents are defined as 

Z e = secondary excitation impedance, Q, 
Z s = secondary winding impedance, ft, 
Z b = burden impedance, ft, 
E s = secondary excitation voltage, ft, 
V t = secondary terminal voltage, V, 

I e = secondary excitation current, A, and 

I s = secondary current, A. 

The burden refers to the impedance of the external load 
applied to the CT secondary, including the impedance of 
the relay, its associated wiring connections, and any 
meters. As shown in figure 10.15, a portion of the gener- 
ated current (I e ) is consumed in exciting the CT core. The 
remainder of the generated current (I s ) is the true value of 
the secondary current. 

The percent ratio correction error (IJIJ is defined as 
that factor by which the nameplate ratio of a CT must be 
multiplied to obtain the true ratio (6). It is apparent that this 
error will remain small as long as the excitation current is 
small. The magnitude of the excitation current is a function 
of the excitation voltage (E s ) as illustrated in the typical 
secondary-excitation characteristics of figure 10.16 (4). It can 
be noted that the secondary-excitation curves are linear until 
the saturation point is reached. Beyond the saturation point, 
a small increase of E s results in a large increase of I e , which 
in turn causes the percent ratio error to increase dramati- 
cally. The magnitude of E s is primarily a function of the 
secondary current (I s ) and the burden impedance (Z b ). There- 
fore, to minimize inaccuracies of the relay system, the 
burden impedance should be kept as low as possible. The 
pickup value for the relay must always result in the excita- 
tion current's lying in the linear portion of the secondary- 
excitation characteristic. It should also be noted that the 
impedance of electromechanical relays is not constant. Since 
they are magnetic devices, their impedance decreases as the 
secondary current increases because of saturation. Thus, the 



Z c X, 




Figure 10.15.— Model of CT and its burden. 



273 



< 



O 
> 



o 
u 

X 

UJ 



ig^ 500/5 
^?-400/5 

300/5 O^/R 




0.005 001 



0.05 0. 



5 10 30 



SECONDARY EXCITING CURRENT A 



Figure 10.16.— Typical set of saturation curves for 600/5 multiratio bushing-type CT. 



relay impedance should be considered over the entire opera- 
tion range of currents when matching the relay to the CT. 

Beyond keeping secondary burden as low as possible, 
there are other recommended guidelines for protective 
relaying that help to minimize the effects of saturation. 
CT turns ratios should be maintained as high as practical. 
Usually, saturation is only a problem on low-ratio current 
transformers, such as 50:5 versus maybe 300:5 ampere- 
turns. CT secondaries with a higher ratio for a specific 
application develop higher voltage and are less likely to be 
saturated under normal burden. Hence the lower the ratio, 
the greater is the chance that a fault will not be cleared 
within the intended zone. Operation of an upstream cir- 
cuit breaker could then cause substantial outages. 

Since underutilizing a CT also produces inaccuracies, 
the maximum anticipated load current should be as close 
to the CT current rating as possible without exceeding it. 
On CT's with 5-A secondaries, IEEE suggests that the CT 
be operated at 3 to 4 A during normal full-load currents 
(4). It has also been suggested that the CT and relay values 
be chosen such that the relay tap setting is at least 
one-half the current rating of the CT secondary (5). If this 
cannot be achieved, the performance of the CT should be 
checked. 

In applications where saturation must be completely 
prevented, the CT should be sized to carry twice the peak 
flux associated with the symmetrical ac fault current (5). 
In most cases, it is not necessary to prevent saturation 
totally in order to provide adequate relaying. For instance, 
immediately after the initiation of a fault, the fault 
current might contain a substantial dc component. This 
component could cause the CT to become saturated, but in 
most mining applications, the dc component decays in less 
than x /i cycle. As high-voltage circuit breaker trip times 
are usually 3 to 4 cycles, the saturation during the first l /z 
cycle would be insignificant in terms of protection. If 



operation is in the linear curve portion, the CT would be 
functioning satisfactorily. Nevertheless, the maximum 
available fault current should be not than 20 times the CT 
current rating, and CT performance should be checked if 
the maximum is above this limit (6). 

Accuracy Calculations 

For a CT under a specific burden, the accuracy of its 
operation (transformer performance) can be calculated 
easily using manufacturer data, provided that the CT has 
a "C" accuracy class. As an example, suppose a 600:5 
multiratio CT is connected through 50 ft of No. 12 AWG 
conductor to an induction-disk relay and an ammeter. 
Both overload and short-circuit protection are intended. 
Figure 10.16 gives the typical set of CT saturation char- 
acteristics, and burden data are contained in table 10.6 (4). 
The primary current (system line conductor) has 24,000-A 
symmetrical rms of available fault current. The objective 
of the calculation procedure is to find the percent ratio 
error. 



Table 10.6. 



-Burdens of relay elements and ammeter 
connected to CT's 



Element 

Relay, timed element, 4-12 A 

pickup. 
Relay, instantaneous element, 

10-40 A pickup. 

Ammeter 

Conductor (interconnecting) 

Transformer secondary resistance... 



Burden 

2.38 VA at 4 A at 0.375 pf; 146 VA 

at 40 A at 0.61 pf. 
4.5 VA at 10 A; 40 VA at 40 A at 

0.20 pf. 

1 .04 VA at 5 A at 0.95 pf. 
0.08 fiat 1.0 pf. 
0.298 Q at 25°C. 



Although the error should not exceed 10% under the 
most extreme circumstances (20 times the secondary cur- 
rent), satisfactory protective-relay operation often calls for 
smaller error, on the order of 2% or less. To compute the 



274 



error, CT secondary burden and relay operating voltage 
must be known. The voltage shown in figure 10.16 is 
related to the CT secondary exciting current, the source of 
the secondary current error. 

Burden should be calculated at maximum secondary 
current. In most cases this is the pickup current for the 
element providing short-circuit protection. Thus, the in- 
stantaneous element of the relay is assumed to be set at 40 
A, which is 8 times rated secondary current or 4,800 A 
primary current. It was stated in chapter 5 that a direct 
summation of the burden voltampere ratings will produce 
adequate results, and while this is generally true, for 
precise work the impedance of each CT load should be 
found and then summed. 

The time-element burden is 146 VA at 40 A and 0.61 
pf, or an impedance, 



Z t = rnlcos 
I 2 



l!pf = :^|52< 



(4or 



Therefore, the percent ratio error correction can be found 
by 



% ratio error = -^ (100), (10.9) 



or at a secondary current, I s , of 40 A, 



% ratio error = -^-(100) = 0.1%. 



Obviously, the accuracy for this application is extremely 
good. 

Only one accuracy calculation example Sas been 
shown because as long as the data are available, this 
technique can be used for any CT application at any 
pickup current. 



or 



Z = 0.091 |53^ = 0.0557 + jO.0723 Q. 

For the instantaneous element, 

S = 40 VA at 40 A 0.20 pf 



or 



Z= = 40 



5 |78^ = 0.005 + jO.025 fi. 



(40) 



For the ammeter, negligible saturation will occur at 8.0 
times rated current (essentially an air-core magnetic cir- 
cuit), and its impedance at 5.0 A is about the same as at 40 

A, 



Z a = 



(1.04) 
(5) 2 



18° = 0.039 + jO.012 U. 



The interconnecting wire resistance (R w ) and CT resis- 
tance (R ct ) are 0.08 and 0.298 fi, respectively. Conse- 
quently, as all burdens are in series, the total burden 
impedance at 40 A is 

Z = Z t + Z ; + Z a + R w + R ct 



= 0.477 + jO.110 = 0.489 | 12.9° Q. 



The magnitude of CT secondary voltage needed to 
produce 40 A is then 

V = IIZI = (40X0.489) = 19.6 V. 



From figure 10.16, the secondary exciting current, I e , at 
this voltage is 

L = 0.03 A. 



LOW-VOLTAGE CIRCUIT BREAKER TRIPS 

In the preceding discussion, overload and short-circuit 
pickup settings were primarily related to external relays 
driven by CT's. The principles of pickup can also be 
applied to low-voltage power circuit breaker settings and 
the instantaneous magnetic settings of molded-case units. 
In this context, pickup would be the minimum current 
within a specified tolerance (usually + 10%) that would 
cause the trip element to activate the operating mecha- 
nisms (4). Although not entirely accurate, pickup can be 
loosely used to describe overload current as specified by 
the thermal trip elements of molded-case circuit breakers. 
However, as element tripping is a function of stored heat, 
tripping times rather than pickup are often employed. 
This section will focus on overload and short-circuit trip 
settings with molded-case breakers and then briefly out- 
line trip levels for low-voltage power circuit breakers. 



Overload Protection 

Overload protection by thermal-magnetic molded-case 
breakers is mentioned in chapter 9. To review, most 
manufacturers calibrate their breakers to carry 100% of 
the continuous current at 40° C; thus, maximum contin- 
uous current is the rating of the thermal-trip unit. Derat- 
ing is only necessary in noncompensating devices when 
the ambient temperature exceeds 40° C. With these 
thoughts, the overload procedure given for relay pickup 
protection can be adapted to this continuous current. The 
thermal element itself defines overload at 125%. For 
instance, the minimum trip time for a typical unit is 
around 10 s for 500% or more of the current rating, with 
time increasing inversely to about 1,800 s at 135% of the 
current rating (20). 

Sections 75.900 and 77.900, 30 CFR (26), define 
breaker use according to Federal law for short-circuit and 
overcurrent (overload) protection of low-voltage and 
medium-voltage power circuits serving three-phase ac 
equipment. No further references are made to overcurrent 
protection except in section 75.900-2(c), which specifies 
that a circuit breaker protecting more than one branch 
circuit must be sized to afford overcurrent protection for 



275 



the smallest conductor. Despite these references to over- 
current protection, Federal regulations are regularly in- 
terpreted and enforced to mean that overload protection is 
not required at each circuit breaker protecting a trailing 
cable (17). 

Perhaps the lack of mandatory cable overload protec- 
tion can be best explained in the following quotation from 
a design engineer working with ac power centers (21). 

Breaker thermal trip ratings should be applied to 
insure complete continuity of operation through 
all overload current peaks rather than to give 
complete cable protection. Many applications are 
being made with no breaker overload protection. 
Short circuit protection only, by means of the 
adjustable magnetic trip units, takes care of line 
to line cable faults. Motor control equipment on 
the machine takes care of motor overloads and 
may give some overload protection to the cable. 

Some States have more stringent requirements than 
the Federal law for overload protection. For example, 
Pennsylvania requires overload protection on all breakers 
serving trailing cables (12). However, there are no specific 
requirements for sizing this overload protection, and so the 
engineer must select an appropriate size for the thermal- 
magnetic circuit breaker. 

Short-Circuit Protection 

The most common short-circuit protection is for trail- 
ing cables and is provided by the magnetic-trip elements of 
molded-case breakers. For surface mines, the pickup re- 
quirements listed for relay pickup protection are used; 
however, for underground coal mines, a set of maximum 
instantaneous settings based on conductor size are man- 
dated in section 75.601-1, 30 CFR (16). These settings, 
listed in chapter 9, table 9.3, apply to both ac and dc 
systems and are based on (2): 

1. An ideal 250-Vdc source feeding a bolted fault at 
the extremity of a 500-ft two-conductor trailing cable, and 

2. A 50% safety factor to allow for circuit breaker 
tolerance, system impedance, and so forth. 

These values are presently in effect for all underground 
coal mine trailing cables. 

Fesak (2) has expressed concern that the dc basis of 
these values does not provide an adequate safety margin 
for resistance-grounded three-phase cables and has recom- 
mended a new set of maximum allowable instantaneous 
settings, which are shown in table 10.7. Here, a specific 
setting is selected not only by conductor size, but also by 
cable length and system voltage. The table stops with 4/0 
cables because this is the maximum practical trailing- 
cable size for underground mines, but the recommenda- 
tions extend to 1,000 MCM. The calculations on which 
these values are based assume an arcing line-to-line fault 
to find the minimum short-circuit current. The multipliers 
to obtain the minimum values from bolted three-phase 
fault currents are 0.85 at 480 V, 0.9 at 600 V, and 0.95 at 
1,040 V. 

Similarly, Vilcheck (18) has recommended a refined 
set of maximum instantaneous settings for dc trailing 



Table 10.7.— Recommended instantaneous trip settings for 
480-, 600-, 1 ,040- V three-phase trailing-cable protection 



Conductor size, 
AWG 



Cable length, 
ft 



Maximum instantaneous 
circuit breaker setting, A 



480 V 



600 V 



1,040 V 



14.. 
12.. 
10.. 
8.... 
6.... 
4.... 

3.... 

2.... 

1.... 

1/0. 

2/0. 

3/0.. 

4/0.. 



0- 500 

0- 500 

0- 500 

0- 500 

0- 550 

0- 500 

501- 600 

0- 500 

501- 650 

0- 500 

501- 600 

601- 700 

0- 500 



501- 
601- 
0- 
501- 
601- 
751- 



600 
750 
500 
600 
750 
800 



0- 500 

501- 600 

601- 750 

751- 850 

0- 500 

501- 600 

601- 750 

751- 900 

0- 500 

501- 600 

601- 750 
751-1,000 



75 

125 

200 

300 

400 

700 

600 

850 

700 

1,000 

900 

750 

1,200 

1,050 

850 

1,400 

1,250 

1,050 

1,000 

1,600 

1,400 

1,200 

1,100 

1,900 

1,700 

1,450 

1,300 

2,050 

1,850 

1,650 

1,350 



100 

150 

250 

400 

550 

850 

750 

1,050 

850 

1,200 

1,050 

950 

1,350 

1,200 

1,050 

1,550 

1,400 

1,200 

1,150 

1,700 

1,550 

1,350 

1,250 

1,950 

1,800 

1,600 

1,450 

2,100 

1,950 

1,750 

1,500 



NAp 
NAp 
NAp 
NAp 
850 
1,200 
1,050 
1,350 
1,150 
1,450 
1,350 
1,250 
1,600 
1,450 
1,350 
1,750 
1,650 
1,500 
1,450 
1,800 
1,750 
1,650 
1,550 
1,900 
1,800 
1,700 
1,600 
1,950 
1,900 
1,800 
1,650 



NAp Not applicable. 



cables, based on minimum expected short-circuit current. 
These pickups are given in table 10.8 and correspond to 
conductor size, cable length, system voltage, and the 
method for grounding. Two types of grounding are in- 
cluded in the table: by means other than a grounding 
conductor, such as diode grounding, and via a grounding 
conductor. Where there is no grounding conductor, only 
line-to-line faults are of concern and the t-i values in table 
10.8 are used. With grounding conductors, line-to-line 
ground faults can also occur, and if no ground-fault protec- 
tion is used, a lower circuit breaker setting is needed for 
cables No. 6 AWG and larger because of the smaller 
grounding conductor ((-g values in table 10.8). 

Low-Voltage Power Circuit Breakers 

Trip settings for low-voltage power circuit breakers 
can involve three direct-acting elements (4). The require- 
ment for long-time delay elements is the same as for the 
overload discussed in relay pickups, and the unit is usu- 
ally set at 100% of the rating, regardless of tolerance. In 
general applications, the short-time-delay element is set 
at five times the overload current point, with the instan- 
taneous element set at nine times the overload level. 
However for mining applications, except for trailing-cable 
protection, the short-circuit requirements covered in relay 
pickups also apply to the instantaneous device. Trailing 
cables demand the same maximum short-circuit settings 
discussed in the preceding section. 



276 



Table 10.8.— Recommended instantaneous trip settings for 
300- and 600-Vdc trailing-cable protection (19) 



Conductor 
size 



Cable 
length, ft 



300-V maximum 

instantaneous 

setting, A 



i-e 



1-9 



600-V maximum 

instantaneous 

setting, A 

t-i t-g 



AWG: 

14 0- 500 50 50 50 50 

12 0- 500 75 75 100 100 

10 0- 500 75 75 200 200 

8 0- 500 100 100 300 300 

6 0- 500 200 100 450 350 

4 0- 500 400 250 650 500 

501- 600 300 150 550 450 

3 0- 500 500 350 700 550 

501- 650 400 250 600 450 

2 0- 500 600 400 800 650 

501- 600 500 350 750 600 

601- 700 450 300 700 550 

1 0- 500 850 600 1,350 1,050 

501- 600 700 500 1,200 900 

601- 750 600 400 1,050 750 

1/0 0- 500 1,050 750 1,550 1,200 

501- 600 900 600 1,400 1,050 

601- 750 750 500 1,200 900 

751- 800 700 450 1,150 850 

2/0 0- 500 1,250 900 1,750 1,400 

501- 600 1,100 750 1,600 1,250 

601- 750 900 600 1,400 1,100 

751- 850 800 550 1,300 1,000 

3/0 0- 500 1,500 1,100 1,950 1,600 

501- 600 1,300 950 1,800 1,450 

601- 750 1,100 750 1,600 1,250 

751- 900 950 650 1,450 1,100 

4/0 0- 500 1,750 1,350 2,150 1,800 

501- 600 1,550 1,150 2,000 1,650 

601- 750 1,350 950 1,800 1,450 

751-1,000 1,050 750 1,550 1,200 
MCM: 

250 0- 500 1,950 1,500 2,300 1,950 

501- 600 1,750 1,300 2,150 1,800 

601- 750 1,500 1,100 1,950 1,600 

751-1,000 1,200 850 1,700 1,350 



f-g Line to ground (with grounding conductor). 
l-l Line to line (without grounding conductor). 



FUSES 



Overload and short-circuit protection can be provided 
by dual-element fuses on low-voltage systems or by power 
fuses on high-voltage systems. The sizing of fuses is 
basically the same as that for relays or tripping devices in 
conjunction with circuit breakers. In cable overload pro- 
tection, the continuous-current rating of a fuse cannot 
exceed the lowest ampacity of any conductor in the pro- 
tected system. Beyond this, as a general protection guide, 
fuses should always be sized to the very smallest 
continuous-current rating that will safely carry any load 
that should be maintained (4). 

Transformer inrush currents, if not heeded, can be a 
problem with high-voltage fuses and for relay pickup as 
well. For instance, upon energizing, a power transformer 
can draw 8.0 to 12 times its full-load current rating for a 
duration up to 0.1 s (4). Expulsion and boric acid fuses can 
be obtained that have inverse-time characteristics to with- 
stand this, thereby affording both overload and short- 
circuit protection for the transformer. High-voltage 
current-limiting fuses, because of their fast response, must 
often have a continuous-current rating to account for 



inrush. In these instances, current-limiting fuses can 
provide only short-circuit protection. 

The use of fuses to provide ground-fault protection is 
more complex. In high-resistance-grounded systems, fuses 
cannot be used to protect against ground faults; in fact, 
their use in series with grounding conductors can create a 
serious personnel hazard if they open. During line-to- 
neutral faults, the current in these systems is limited to a 
level lower than typical load current in the line conduc- 
tors; line-to-line-to-neutral faults cause about the same 
current levels. As there are no neutral-connected loads in 
these systems, fault-current contribution is only from 
generating stations and capacitance; that is, motors have 
no effect. 

However in ungrounded or solidly grounded systems, 
ground-fault currents can be substantially greater; thus, 
line-interrupting devices including fuses can provide 
ground-fault protection. For instance, in an ungrounded 
system, fuses sized to the short-circuit requirements dis- 
cussed in this chapter would probably cause interruption 
during two simultaneous line-to-neutral faults on different 
power conductors. 



COORDINATION 

The basic principles of coordination have already been 
discussed in chapter 9 and earlier in this chapter, but here 
the specific procedures used in coordination studies will be 
demonstrated. References 1 -2 are highly recommended for 
more detailed coverage of this topic. 

A coordination study is basically a comparison of the 
times-to-operate for individual protection devices under 
both normal and abnormal current flows (3). A prelimi- 
nary study should always be made in the early stages of 
mine power-system design, since it could indicate that 
transformer sizes or impedances require modification or 
that cable sizes need to be changed. After final selection of 
devices and components, a second study must be made to 
confirm the tentative study. A new coordination investiga- 
tion is required whenever the original system is changed, 
whenever new loads are added, or whenever existing 
equipment is replaced with higher rated components. An 
additional study should always be made if a fault causes a 
major shutdown of an existing system. 

Coordination has two conflicting objectives: protection 
and selectivity (3). Each relay, tripping unit, or device is 
designed to protect a specific circuit segment and for 
maximum safety must isolate each faulted circuit as fast 
as possible. At the same time, continuity of service must 
be provided, and this requires selectivity. Again, selectiv- 
ity is the concept of isolating only that portion of the 
system experiencing failure. This demands successively 
longer fault-clearing times or higher pickup settings for 
each protection zone from the loads to the source. It is the 
role of the coordination study to optimize these require- 
ments without compromising personnel safety. 

A coordination-curve plot is one of the best methods 
for achieving these objectives. For resistance-grounded 
systems, two plots can be made, one for line-conductor 
protection and the other for ground faults. The protection 
described here refers mainly to line-conductor protection 
plots but also applies to grounding. 

Before a curve plot can be constructed, a good one-line 
diagram of the system must be prepared. In addition to the 



277 



normal diagram features discussed in chapter 4, it should 
include the following information {3-4): 

1. Maximum and minimum levels of short-circuit 
current that are expected to flow through each protective 
device (obtained from a system fault analysis); 

2. Normal and maximum load currents for each sys- 
tem portion (from a load-flow analysis); 

3. CT ratios; and 

4. Relay, circuit breaker, and fuse ratings and adjust- 
ment ranges (time-current characteristics must also be 
known). 

With these data available, as shown in figure 10.17, the 
coordination study can proceed to plotting. 

Coordination curve plots are normally drawn on stan- 
dard log-log graph paper, using the vertical scale for time 
and the horizontal axis for a common current source (3). 
For good results, this scale usually corresponds to currents 
at the lowest voltage level, while currents at higher 
voltage levels are plotted on the same scale as equivalents. 
For instance, if utilization is at 600 V with distribution at 
7,200 V, the current scale is calibrated for 600 V. Currents 



in those system portions are plotted directly, but distribu- 
tion currents are multipled by 7,200/600 or 12, then 
plotted. However, note that this is not a firm rule, as any 
convenient scale can be used. Important current levels 
such as transformer inrush, motor starting, and fault 
current can be placed on the graph as points. 

A critical protection path is then chosen from the 
one-line drawing by selecting a critical load or safety 
hazard and following the protection path back to the 
source. Fortunately, because of radial distribution, this 
selection is an easy matter in most mine power systems. 
The time-current characteristics of all protective devices 
along the path or affecting the path are then plotted on the 
graph. 

As shown in figure 10.18 (4), such a plot is invaluable 
as it allows protection-device action to be visualized 
throughout the system on the same current basis. This is 
an advantage similar to that of per-unit analysis. It also 
provides for the normalization of the characteristics for 
each protection device, which rarely have the same shape 
or even scale. The plot simplifies the selection of current 
and time settings that will provide the best possible 
protection and safety while also giving selectivity. One 
way to do this is by specifying the characteristics of the 
most downstream protection, then sequentially position- 
ing each protection step toward the source (or vice versa). 

As is evident in figure 10.18, operating time intervals 
must be maintained between protective devices in order to 



100 E 34.5-kVfuse 

34.4-kV <^_Lkj A 3,750-kVA,Z = 6% 
4,160-V ^ry-^ _£_ 



800/5 




300/5 } (tt 



100-A 
) molded-case 
circuit breaker 




40 100 400 

CURRENT, A 

at 4,160-V multiply by 10 
" 480 -V " " 87 
" 34.5-kV " " 1.21 



4,000 10,000 



Figure 10.17.— Example of one-line diagram for preparing a 
coordination curve plot for one path. 



Figure 10.18.— Coordination curve plot for figure 10.17 
showing various protective-device characteristics. 



278 



achieve correct sequential operation. When applied to 
relay-activated circuit breakers, this time margin must 
allow for circuit breaker interrupting time, overtravel 
time, and a safety factor (4). All these times are additive. 

The time required by a circuit breaker to interrupt 
current once it receives the trip signal is equal to its speed 
in cycles divided by 60. A typical power circuit breaker has 
an operative speed of 5 cycles, corresponding to an 0.08-s 
operating time. Overtravel is important in electromechan- 
ical time-delay relays, especially inverse-time, induction- 
disk types. Owing to moving-part inertia, these relays will 
continue to close their contacts even after the fault is 
removed, and overtravel can thus create a relay operation 
delay. Typical overtravel time for inverse-time, induction- 
disk relays is 0.10 s. The safety factor mainly allows for 
variations in relay characteristics due to manufacturing 
differences, tolerance, aging, and dust; the commonly 
assigned range is 0.12 to 0.22 s. Summing these factors 
results in a necessary time margin range of 0.3 to 0.4 s (4). 
While the typical case is considered to be 0.4 s, the margin 
can be reduced to 0.3 s for carefully tested systems. Note 
that static relays eliminate overtravel and allow the use of 
the minimum safety factor, so that the typical time margin 
can be reduced to 0.2 s (see chapter 14). 

The typical time interval of 0.4 s is used for relay- 
to-relay coordination in figure 10.18, and the time selec- 
tions for line-conductor relays can be seen. Although this 
time interval considered only external relay tripping, the 
0.4-s time margin is often utilized even when direct-acting 
low-voltage circuit breakers are coordinated together or 
with relayed interrupters. The same is true when a circuit 
breaker is coordinated with an upstream fuse if the total 
fuse clearing time is greater than 1.0 s. With less than 1.0 
s, the fuse-to-circuit-breaker time margins can be reduced 
as low as 0.1 s (4). 

For an additional example of coordination, consider 
ground-fault protection in a high-resistance-grounded dis- 
tribution system. A separate plot can be made for the 
ground-fault conditions. If the induction-disk relay at the 
most downstream switchhouse is set to pick up with the 
minimum time dial setting, the time dial on the ground- 
fault relay in the next upstream switchhouse would be set 
0.4 s higher. At each subsequent upstream switchhouse, 
the ground relay time would be set progressively higher 
until the neutral bushing of the source transformer is 
reached. These time allowances also involve the current 
equivalent of any potential relaying about the grounding 
resistor. Because of the need to maintain the time inter- 
vals, many engineers have found the coordination of more 
than five relays in one ground-fault current path nearly 
impossible, since the maximum practical time setting is 
about 2.0 s. Although coordination curve plotting has been 
discussed here, in practice the manufacturer characteris- 
tic curve could be used in this case. The reason is that 
available ground-fault current in a high-resistance- 
grounded mine distribution system remains fairly con- 
stant, so that the same very inverse relays could be used 
for all ground-fault protection, with the possible exception 
of resistor potential relays. 

The material presented in this chapter and the last 
has perhaps given the impression that in order to create a 
safe mine power system all circuits should have overload, 
short-circuit, and ground-fault protection. While this is 
generally true, it is not quite the entire picture. Worker 
safety must always come first; hence, equipment protec- 
tion is not applied in those cases where its use can cause 



hazards to personnel. These exceptions almost always 
involve unit-designed equipment, such as elevator motors 
or the swing, hoist, and propel motors of surface excava- 
tors. The only other exceptions to mandatory equipment 
protection are cases where there is no possibility of com- 
promising personnel safety if it is omitted. 

This chapter has introduced the important parame- 
ters used in the selection and adjustment of protective 
devices. This information together with the material in 
chapter 9 serves as one basis for chapters 12 and 13, where 
the components and techniques are applied to the equip- 
ment used in the mine system. But first another type of 
protection must be discussed, protection against the some- 
times hidden dangers of transients and overvoltages. 



REFERENCES 

1. Crites, W. R., and A. G. Darling. Short-Circuit Calculating 
Procedure for D-C Systems With Motors and Generators. Trans. 
Am. Inst. Electr. Eng., Part 3, v. 73, Aug. 1954. 

2. Fesak, G., W. Helfrick, W. Vilcheck, and D. Deutsch. Instan- 
taneous Circuit Breaker Settings for the Short Circuit Protection 
of Three Phase 480, 600 and 1040 V Trailing Cables. Paper in Con- 
ference Record -IAS 12th Annual Meeting (Los Angeles, CA, Oct. 
1977). IEEE, 1977. 

3. Institute of Electrical and Electronics Engineers (New York). 
Recommended Practice for Electrical Power Distribution for In- 
dustrial Plants. Stand. 141-1986. 

4. Recommended Practice for Protection and Coordina- 
tion of Industrial and Commercial Power Systems. Stand. 
242-1986. 

5. Kiefer, J. A., and J. L. Kohler. Ground Fault and Overcurrent 
Protection Criteria for Coal Mine AC Distribution Systems (con- 
tract J0395035, Ketron, Inc.). BuMines OFR 158-81, 1980; NTIS 
PB 82-137001. 

6. Mason, C. R. The Art and Science of Protective Relaying. 
Wiley, 1956. 

7. Morley, L. A., F. C. Trutt, and R. A. Rivell. Coal Mine Elec- 
trical System Evaluation. APL Mine Electrical System Load-Flow 
Program (grant G0155003, PA State Univ.). BuMines OFR 
61(7)-78, 1977; NTIS PB 283 496. 

8. Coal Mine Electrical System Evaluation. Extended 

APL Mine Electrical System Load-Flow Program (grant 
G0155003, PA State Univ.). BuMines OFR 55-78, 1977; NTIS PB 
283 497. 

9. Myers, W. P. Current-Limited Ground Fault Relaying. Min. 
Congr. J., v. 59, Apr. 1970. 

10. National Fire Protection Association (Quincy, MA). National 
Electrical Code. NFPA 70-1981 (ANSI Cl-1981). (Updated every 3 

yr.) 

11. Paice, D. A., and A. B.. Shimp. Discriminating Protection 
for Trolley Wires. Paper in Mine Power Distribution. Proceedings: 
Bureau of Mines Technology Transfer Seminar, Pittsburgh, Pa., 
March 19, 1975. BuMines IC 8694, 1975. 

12. Pennsylvania Department of Environmental Resources. 
Bituminous Coal Laws of Pennyslvania, Sec. 333. 1961. 

13. Poker, L. R. Economical Ground Fault Protection Available 
With a Standard Low-Voltage Tripping System. IEEE Trans. Ind. 
and Gen. Appl., v. 6, Mar./Apr. 1970. 

14. Shimp, A. B., and D. A. Paice. Application of Molded-Case 
Breakers on DC Electrical Systems in Coal Mines. Paper in Mine 
Power Distribution. Proceedings: Bureau of Mines Technology 
Transfer Seminar, Pittsburgh, Pa., March 19, 1975. BuMines IC 
8694, 1975. 

15. Stagg, G. W., and A. H. El-Abiad. Computer Methods in 
Power System Analysis. Mc-Graw-Hill, 1968. 

16. U.S. Code of Federal Regulations. Title 30-Mineral 
Resources; Chapter I -Mine Safety and Health Administration, 



279 



Department of Labor; Subchapter O-Coal Mine Health and Safe- 
ty; Part 18 -Electric Motor-Driven Mine Equipment and Ac- 
cessories; Part 75 -Mandatory Safety Standards, Underground 
Coal Mines; Part 77 -Mandatory Safety Standards, Surface Coal 
Mines and Surface Work Areas of Underground Coal Mines; 1981. 

17. U.S. Mine Safety and Health Administration. Coal Mine 
Safety Electrical Inspection Manual, Underground Coal Mines. 
Apr. 1979. 

18. Vilcheck, W., G. Fesak, and W. Helfrich. Instantaneous Cir- 
cuit Breaker Settings for the Short-Circuit Protection of Direct 



Current 300 and 600 Volt Trailing Cable. Paper in Conference 
Record -LA.S 13th Annual Meeting (Toronto, Ontario, Canada, Oct. 
1978). IEEE, 1978. 

19. Wade, E. C. Ground Relaying for Mining Distribution 
Systems. Coal Age, v. 71, July 1966. 

20. Westinghouse Electric Corp., Low-Voltage Breaker Div. 
(Beaver, PA). Breaker Basics. 1973. 

21. Youel, V. H. The Underground A-C Mine Power Center. 
Mechanization, v. 24, Aug. 1960. 



280 



CHAPTER 1 1 .—TRANSIENTS AND OVERVOLTAGES 



The term electrical transient can mean different 
things to people of different interests. However, there are 
some common ideas. As defined by Greenwood (19), 1 "an 
electrical transient is the outward manifestation of a 
sudden change in circuit conditions, as when a switch 
opens or closes, or a fault occurs on a system." The circuit 
parameters of inductance and capacitance are found in 
any power system to some extent. When the system is 
changed, the quantities of current, voltage, magnetic flux, 
and so on, do not instantly assume new values. Rather, 
they go through a transition to reach the new steady-state 
condition (33). It is this transition period that gives rise to 
the transient voltages and currents. 

Transients are of short duration, and the time in 
which they occur is an extremely small percentage of the 
total operating time. But it is during these short periods 
that some of the greatest electrical stresses can occur, 
mainly because of excessive currents or voltages. The 
excessive voltage is often critical in the design of mine 
power systems. In extreme cases, system parts are dam- 
aged and equipment failure follows. Safety can be further 
compromised because anomalous voltages may exist at 
machine frames. 

Transients are a fact of life on every power system, yet 
careful design within the technical and economic con- 
straints of a system can result in a reduction of transient- 
related component failures. However, mine power systems 
are particularly vulnerable to transient-induced failures 
because their operation and arrangements are extremely 
dynamic. As a mining activity advances, the electrical 
system is expanded, often on a weekly basis. Although this 
expansion is normally designed into the system, circum- 
stances in the mine can call for additional modifications. 
In some cases, these on-the-spot modifications may not be 
in line with sound engineering principles. 

A lack of knowledge about transients has resulted in 
power-system requirements that could be misunderstood 
by some mine power engineers. Furthermore, in the pro- 
cess of maintaining the system, faulty components are 
sometimes replaced with devices having different specifi- 
cations or even total incompatibility. These two factors 
also increase the vulnerability of mine power systems to 
transient occurrences. 



TRANSIENT SOURCES 

There are several discrete sources of transient over- 
voltages. Although slight differences exist in classifying 
these events, IEEE standards catalog seven types (23): 



references 19, 33, and 37 are invaluable sources of tran- 
sient information, and these should be consulted for de- 
tailed coverage of transient phenomena and problems. 



LIGHTNING PHENOMENA 

On a global scale, the earth and the lower part of the 
ionosphere may be considered as conductive bodies sepa- 
rated by a rather poor conductor, the atmosphere. The 
entire system is analogous to an enormous capacitor with 
a leaky dielectric. Any charge unbalances that accumu- 
late are dissipated by sudden breakdowns in the dielectric 
(atmosphere), and the resulting current (lightning) is of 
short duration but high amplitude. At any given time, 
electrical storms are taking place somewhere on earth, 
and the average current flow between air and earth is 
more or less constant at a level of 1,500 A. During fair 
weather, the normal voltage gradient in the air near the 
earth's surface is about 3.0 V/cm, but this rises to around 
500 V/cm beneath a thundercloud. The potential differ- 
ence needed to initiate a lightning stroke is on the order of 
50 million V (12). In the continental United States, a 
typical stroke lasts for only a fraction of a second, yet 
releases a tremendous amount of energy, approximately 
200 million J. The total quantity of charge involved in the 
average stroke is about 20 C, and the peak current value is 
around 20 kA (28). 

As shown in figure 11.1, the shape of the discharge 
starts with an extremely short-duration voltage pulse, 
whose crest may reach 5 MV or more (29). This is closely 
followed by the current waveform, which rises more slowly 
to its peak value and lasts longer. About 82% of the strokes 
occurring in the United States are negative in polarity; 
that is, they transfer electrons from the clouds to earth. 
Figure 11.2 illustrates the wide distribution of current 
magnitudes that may be expected in various lightning 
strokes (27). 

Figure 11.3 shows the expected number of thunder- 
storm days per year for any of the 48 contiguous States 
(27). The probability that a particular object will be hit by 
lightning depends upon the cloud charge intensity, the 
geographical locale, and the height of the object in relation 
to its surroundings. Charge intensity is a parameter 
because an increased charge in the tip of the stroke 
generates a proportionally higher electric-field gradient. 
This determines the attractive range of the stroke, in 
other words, the horizontal distance from the tip of a 
downcoming leader to an object receiving the stroke. As 



1. Lightning, 

2. Switching surges, 

3. Static, 

4. Contact with a higher voltage system, 

5. Line-to-ground faults, 

6. Restriking ground faults, and 

7. Resonant conditions. 

The following paragraphs introduce these sources briefly 
in light of mining applications. It should be noted that 



1 Italicized numbers in parentheses refer to items in the list of references 
at the end of this chapter. 



VOLTAGE : Rise rate from 1 to 3 MV///S 
/ Duration on the order of a few microseconds 



HIGH CURRENT : Maximum value about 100 kA 
Duration, lOtolOO^s 



/ 




LOW CURRENT : Value, 100 to a few thousand amperes 
\ Duration, 0.001 to 0.01 s 



II I I 

Figure 11.1.— Schematic representation of lightning stroke 
discharge. 



281 



200 - 



160 - 



LU 
01 

en 



LU 

cc 
o 




0.02 0.1 0.2 12 5 10 20 50 100 

STROKES WHERE CURRENT EXCEEDS ORDINATE, % 

Figure 11.2.— Distribution of crest currents in lightning 
strokes. 



shown in figure 11.4, the striking distance for a typical 
20-kA stroke is about 30 m or 100 ft (27). 

The exposed surface portions of a mine electrical 
distribution system are particularly vulnerable to light- 
ning. For instance, overhead lines on wooden poles are 
often used to carry three-phase power from substations to 
surface facilities of underground mines. In surface mines, 
the overhead lines can extend close to the shovels and 
other electrically operated equipment in the pit. Since 
these poles can also carry the grounding conductor from 
the safety ground bed, it is doubly important that they be 
adequately protected from lightning. 

Three separate failure modes are recognized whereby 
an electrical distribution system can be damaged by 
lightning. If the impedance to ground of a supporting 
structure is high and the structure is struck by lightning, 
its potential may become elevated above ground to the 
point that a flashover occurs from the tower to a power 
conductor. This is known as backflashover. If a high- 
intensity stroke to earth occurs nearby, the resulting 
electric field may be great enough to induce excessive 
potentials on conductors; a power conductor can then arc 
over to the tower. In a shielding failure, lightning hits a 
power conductor directly, raising its potential to the point 
where arcing takes place between the conductor and the 
support structure (27). Induced potentials on grounding 
conductors, including equipment frames, can elevate the 
potential of the grounding system above that of earth and 
cause damage to protective devices or create a personnel 
hazard. With nearby strokes to earth, the magnitude of 
the induced voltage depends upon stroke current, distance, 
and height of the conductor, as illustrated in figure 11.5 
(25). 



SWITCHING TRANSIENTS 

Switching operations account for the majority of all 
transient phenomena on mine power systems, and except 
for direct strokes by lightning, the most destructive power- 
system transients are initiated by this source. Every time 
a switch is operated, a transient can occur, and these 
transients can be classified into two types: normal and 




Figure 11.3.— Map showing average number of thunder- 
storm days per year in the United States. 



240 - 




Positive 



Negative 



LU 

o 

< 

I- 
co 

a 
o 

\- 



40 80 120 

LIGHTING-CURRENT AMPLITUDE, kA 

Figure 11.4.— Striking distances for negative and positive 
strokes. 



DISTANCE A FROM STROKE CHANNEL, ft 
1,000 600 400 200 150 100 75 




200 400 600 
400 800 1,200 
800 1,600 

CREST VOLTAGE, kV 



800 



1,000 /?=25ft 
/?-50ft 
/?= 100ft 



Figure 11.5.— Crest voltages induced on transmission lines 
by nearby strokes. 



282 



abnormal (33). Normal voltage and current transients are 
considered a characteristic of proper operation, and their 
maximum outcome is a transient voltage or current no 
greater (but usually less) than two times the peak steady- 
state voltage or current of the system. Transients exceed- 
ing this level are termed abnormal, and these are the 
result of incorrect component operation or design. Abnor- 
mal switching transients can occur in several ways, but all 
involve the release or exchange of trapped energy in the 
power-system inductance or capacitance (19, 33). Interest- 
ingly, a normal transient can be generated on a quiet 
power system, but if energy from this transient is stored 
and afterwards a second transient is created, the latter can 
be abnormal. 

Transient phenomena caused by switching can be 
divided into those created on ac systems and those associ- 
ated with dc. The ac switching events can be further 
subdivided into capacitive switching, current chopping, 
and prestrike. 

Capacitance Switching 

Because of the extensive use of cables, surge capaci- 
tors, and power-factor correction capacitors, capacitance 
plays a major role in most mine electrical systems. Abnor- 
mal transients can occur when ac current to capacitance is 
interrupted (6, 19). 2 The problem is that current leads 
voltage by nearly 90°. Capacitive currents are usually 
rather small so there is a good chance that current flow is 
stopped at a very early current zero. Thus, the load-side 
capacitance would be charged to the peak voltage of the 
system, and this energy becomes trapped. 

A simple circuit to demonstrate capacitance switching 
is shown in figure 11.6, where an ac source is feeding cable 
or line lumped capacitance through a circuit breaker. 
Source voltage and current waveforms are given in figure 
11.7A; the point of current interruption is signified by the 
dashed vertical line. Figure 11. IB shows the voltage 
across the capacitance, and figure 11. 7C relates the poten- 
tial or recovery voltage across the circuit breaker contacts, 
which is the difference between the source and load sides. 
System inductance and stray capacitance are neglected in 
these waveforms to illustrate the effects of the trapped 
charge (19). 

Because early interruption is possible, the recovery 
voltage may attain two times the system peak when the 
contact gap is quite small. Hence the arc may reignite or 
restrike. With restrike, an inductance-capacitance circuit 
is formed with a resonant frequency of 



Source 
L 



© 



Breaker 



' l J- Lumped 
capacitance 



Figure 11.6.— Simple circuit to illustrate capacitance- 
switching voltage transients. 






A Source -side 
voltage 



B Load -side 
voltage 




C Recovery voltage 



Figure 11.7.— Voltage and current waveforms before and 
after current interruption. 



The basic equation for the per-phase voltage after restrike 
is 



f„ = 



(11.1) 



T dl 



(11.2a) 



where f = resonant or natural frequency of system, Hz, 

L = system inductance, H, 
and C = system capacitance, F. 



2 Personal communication from E. K. Stanek, West Virginia University, 
Aug. 1977. 



where v = V m cos(wt) = source voltage, V, 
V m = peak source voltage, V, 

v c = V c(0) + p \ idt = capacitance voltage, V, 
V c(o) = voltage across capacitance at restrike, V, 
and i = current through circuit breaker, A, 



or 



di 



L dt + h J idt = Vm cos (wt) " v °- (11 - 26) 



283 



lb investigate the current transient in the short time 
interval after restrike, 



and thus, 



cos(wt) ~ 1, 



L oi + ^ idt = v ~- v °- 



(11.2c) 



Solving this equation for the time-domain current, it is 
found that 



i(t) = (V m - V ) (£) ' sin ( Wo t). 



(11.3) 



Voltage 



Current 



60-Hz current 



approaches 




Figure 11.8.— Voltage and current transient waveforms oc- 
curring with capacitance switching and restrike. 



Accordingly, the voltage across the capacitance would be 

1 ft /C\ - 5 

v c = V + -j (V m - V )(-) Bin(« t)dt 



or 



or 



v. = 



V - V ft 
= V Q+ ™ C)0 .5° ) sin(o> t)dt 



v c = V + V ^ C)0 J° [1 - cos (co t)]. (11.4) 



1.0 mH 



2,300 V 



Contactors 




25-ft SHD cable 



/ 



100 kvar 
(50^F) 



h 

(\ 500-hp 



Figure 11.9.— Per-phase diagram of 4,160-V pump-motor cir- 
cuit. 



For the worst case, 

V = -V 

* o * m 

or 

V - V = 2V , 

T m o m> 

and the transient current is 

i(t) = 2V m Q a5 sin (« t), 



(11.5a) 



with transient voltage per phase for the load-side capaci- 
tance being 



v c = V m - 2V m cos ( Uo t). 



(11.56) 



Both these sinusoidal waveforms are illustrated in figure 
11.8; the oscillations are at the system natural frequency 
(19). Equation 11.56 is of special interest as it indicates 
that the transient voltage can reach three times the peak 
system voltage. The decay shown in the waveforms is due 
to damping by system resistance. Obviously, system resis- 
tance is neglected in the foregoing derivation, but its effect 
is considered small compared with capacitance and induc- 
tance in this type of switching. 

To show the significance of the current transient, 
consider figure 11.9, which is a per-phase diagram of a 
three-phase motor circuit. Here, three-phase 4,160-V 
power is being fed through a circuit breaker (used as a 
motor starter) through 25 ft of SHD cable to a 1, 500-hp 
wound-rotor motor. A 300-kvar capacitor bank is used for 
power-factor correction and is placed at the load side of the 



breaker. Assume that the motor is drawing very little 
current, the breaker is opened, and interruption occurs at 
the first current zero. If the motor current is very small 
compared with the current drawn by the capacitor, the 
stage is set for restrike. Peak restrike current would be 
approximately 



i P = 2 Vp g»-= 



or 



Ip = 2(3,396)i 



5 x 1Q- 

10~ 3 



= 1,520 A. 



Since the steady-state current through the capacitor is 
around 45 A, this transient-current peak is about 34 times 
the normal capacitance current flow, approaching a level 
expected for motor-starting inrush. 

Even though this is significant, the serious problem is 
not the current transient but the possible overvoltage 
produced. It can be seen from equation 11.56 that, after 
one restrike, the per-phase voltage can approach 3 V m , 
3(3,396) V, or about 10 kV. This is substantial but still 
within the insulation capabilities of most 4,160-V systems. 
However, if the capacitor voltage is still around 3 V m when 
current becomes zero again, a current interruption will 
trap 3 V m across the capacitor. Now, the circuit breaker 
recovery voltage may approach 4 V m , causing a second 
restrike, which can then cause a capacitor voltage of - 5 
V m . This situation can continue, developing even higher 
voltages. Such a multiple-restrike process is shown graph- 
ically in figure 11.10 (33). The voltages created can cause 
insulators to flashover. 



284 



The problem here is large capacitance as seen by the 
switching apparatus, with a very small impedance be- 
tween that capacitance and the load-side contact. Exam- 
ples of this situation could include the charging current to 
an unloaded cable or distribution line. Yet perhaps the 
most drastic instance would be interrupting current to a 
static capacitor bank, involving one line of a grounded 
bank or two lines of an ungrounded bank. 

Current Chopping 

Current chopping is the phenomenon of forcing cur- 
rent to zero before a natural current zero. This can occur 
when small currents, such as transformer magnetizing 
current, are interrupted by switching apparatus, as illus- 
trated graphically in figure 11.11 (19, 26). Current chop- 
ping can trap magnetic energy in the power-system seg- 
ment being interrupted, and the result can be severe 
transient voltages. 

In recent years, current-chopping transients have re- 
ceived more attention than any other type. The concern 
has been connected with the increased use of vacuum 
circuit breakers (VCB's), perhaps to the point where these 
transients have become associated with the use of VCB's. 
Through their high efficiency, these interruptors can eas- 
ily chop small current flow. However, it should be noted 
that all circuit breakers can cause current chopping, as 
well as certain fuses, especially the current-limiting types. 

To appreciate the magnitude of the created overvolt- 
ages, consider the simplified equivalent circuit of a power- 
system segment shown in figure 11.12. The components 
can be considered per phase for the three-phase system. If 
the circuit breaker chops the current at magnitude I, 
magnetic energy is stored at that instant in the system 
inductance (mainly the transformer) at a level (33) 



W L = 



LI 2 



(11.6) 



where W L = energy stored, J, 

L = transformer magnetizing and cable induc- 
tance, H, 
and I = magnitude of chopped current, A. 

This stored energy is then transferred to the capacitance 
and charges the capacitance to 



W„ = 



CV 2 



(11.7) 



where V = voltage produced across capacitance, V, 
and C = lumped capacitance of cable and transformer, 
F. 

Neglecting losses, both energies are equal, producing a 
voltage 



Source voltage 



Source voltage 
and current 



Capacitor voltage 







Recovery voltage 



Figure 11.10.— Voltage and current waveforms resulting 
from multiple restrikes after capacitance switching. 




Figure 11.11.— Graphic example of current chopping by 
breaker interruption. 



-• •- 



Transformer 
and cable 



Source r\; 



Circuit 
breaker 
contacts 



Ts' 



Figure 11.12.— Equivalent circuit of power-system segment 
with lumped components per phase, neglecting resistance. 



Vp-I 



= iz„ 



(11.8) 



ferral back and forth causes oscillations with a frequency 
equal to 



where V p = peak transient voltage, V, 
and Z = system surge (or characteristic) impedance, 
U. 

Following the capacitance charging, the energy is trans- 
ferred back to the system inductance. The effect of trans- 



Ry 



f = i[J__(A) 

° 2tt LLC \2L/ 



(11.9) 



where f = oscillation frequency, Hz, 
and R = system resistance, 12. 



285 



Equation 11.9 shows the damping effect of system resistance, 
but for very small resistances (as are usually the case in 
power systems) the equation reduces to approximately 



f„ = 



(11.10) 



This is again the natural frequency of the power system. 
The preceding has considered the theoretically pure 
world and, of course, actual systems exhibit losses. 
Present-day dry-transformer hysteresis losses limit energy 
storage to about 40% (20). Therefore, the peak transient 
voltage is restricted to about 63% of equation 11.8 or 



V p = 0.631 



^ = 0.613IZ o 



(11.11) 



0.63 IZ n - 




Transient recovery voltage 



Transient waveform 



TIME 



/ 
s 

■-- < 

Normal waveform if 
no interruption 

Figure 11.13.— Graphic example of chopping voltage tran- 
sients. 



An illustration of a voltage transient resulting from chop- 
ping (by trapping energy) is given in figure 11.13. 

Consider the power system shown in figure 11.14, 
where a switchhouse is connected to a load center by a 
short cable. A typical switching procedure in mining is to 
interrupt or switch out an unloaded load center, for 
example, during a maintenance shift. Even though there 
are no connected loads on the secondary, the transformer 
still draws a small amount of magnetizing current, about 
0.03 to 0.05 pu of rated current for most mining applica- 
tions. Assume that the cable is so short that it has 
negligible capacitance and inductance as compared with 
the transformer. Thus, the equivalent circuit would be as 
shown in figure 11.12. 

A typical load center in underground coal mining is 
750 kVA, and common unloaded transformer values for a 
7,200-Vac system are C = 3,000 pF and L = 15 H. If 1.0 A 
is chopped by the circuit breaker, the per-phase crest 
voltage that can be produced on the system is 



<sTIo- 



= 44.5 kV. 



This level is indeed high, being directly proportional to the 
product of the system surge impedance and the chopped 
current. Hence, if 2.0 A is chopped (for example, the peak 
due to distortion), V p would equal 89.1 kV. Significantly, 
the level is independent of rated system voltage but is 
inversely proportional to the square root of the load-side 
capacitance. Common mine power-distribution systems 
employ 4,160, 7,200, or 12,470 V, and transformers of the 
same capacity (kVA) have similar cores, surge impedances, 
and magnetizing currents; therefore, similar energy levels 
can be stored by chopping, and similar peak transient 
voltages produced, regardless of system voltage. 

The magnitude of chopping obviously depends upon 
the instant of switching. For example, if switching oc- 
curred at the natural current zero, no chopping could 
result, and the transient would be normal. Yet the proba- 
bility of this happening is extremely small. Consequently, 
if the surge impedance is high, there is a likelihood that 
serious transients can be created. Configurations similar 
to that described above can also result in large chopping 
transients, for example, a cable-connected motor. However, 
surge impedance and magnetizing current are smaller in a 
motor than in a transformer, which reduces the danger of 
large chopping transients. 



Incoming power 
SWITCHHOUSE 






LOAD 
CENTER 
















/-> 


750 kVA 
7,200 Vac 


No loads 




' Circuit 


Short I/O SHD- 
cable 


■GC 


connected 






breaker 






Fe 


edtf 


trough 





Figure 11.14.— Segment of mine power system. 



In addition to system losses, there are other phenom- 
ena that can reduce the transient voltage. After chopping, 
with only a small VCB contact separation, the dielectric in 
the circuit breaker is often unable to support the transient 
voltage, and an arc restrikes (19). Sometimes the circuit 
breaker will make successive attempts to clear the circuit 
in this manner, reaching progressively higher voltages as 
the contact gap increases, until isolation is finally estab- 
lished. But the maximum voltage attained may not be as 
great as when switching is clean; the stored energy is not 
allowed to accumulate, and the effect can reduce V p by as 
much as one-half (20). Nevertheless, the sequential re- 
striking and clearing of a circuit breaker can create 
dangerous overvoltages, and the voltage escalation can be 
much more rapid (19). 

If the cable length between the switchhouse and the 
load center is increased, the cable capacitance will be 
proportionally increased. This will reduce the surge im- 
pedance and therefore the possible severity of the chopping 
transient. Yet too large a capacitance can result in other 
transient phenomena, not only from restrikes as previ- 
ously covered, but also from prestrikes. 

Prestrike 

During the initial energization of system capacitance, 
an arc ignition or prestriking may occur across the contacts 
of a circuit breaker, prior to (final) mechanical closure (32). 
This phenomenon appears to be enhanced when the load- 
side capacitance is considerably in excess of the source-side 
capacitance. These events seem dependent upon the capaci- 
tive inrush current, and the magnitude of inrush current is 
mainly controlled by the amount of load-side capacitance. 



286 



Prestrike transients have usually been associated with vac- 
uum interrupters, and overvoltages have been found to reach 
substantial levels if the initial inrush current caused by the 
prestrike is momentarily stopped, then followed by a subse- 
quent prestrike or contact closure. Situations can involve 
energizing one line of grounded capacitance or two lines of 
ungrounded capacitance. 

Hypotheses that have been advanced to account for 
prestriking include the following (10): 

• Contact "whiskers." When exposed to a high elec- 
trostatic stress in a vacuum, extremely fine filaments can 
grow from the surface of metals. Such filaments could 
cause an arc ignition. 

• High electric-field strength. Stress created by a high 
electric-field strength could cause breakdown of the dielec- 
tric. 

• Contact bounce. Contact bounce is the deflection of 
the moving contact after impact with the fixed contact of 
the interrupter. This is caused primarily by a weak oper- 
ating mechanism, and the resulting separation could 
allow arc extinction. 

Whatever the source, if the load-side capacitance is 
uncharged and a prestrike occurs, the voltage on the 
immediate load side of the breaker will collapse to zero. 
This can cause voltage and current traveling waves to 
radiate on cable and lines. As will be discussed later, the 
waves can reflect and refract at discontinuities in the 
system characteristic impedance, and the typical frequen- 
cies of the resultant waveforms can be 60, 600, 6,000, 
60,000 and 600,000 Hz, all superimposed. The combina- 
tion of the harmonics obviously has many current zeros, 
and when these zeros occur between the prestrike and the 
mechanical contact closing, the circuit breaker can easily 
interrupt the inrush-current flow. In the same manner as 
in capacitance switching, energy can be trapped on the 
capacitance being energized. Subsequent arc reignitions 
followed by interruptions can create an unusually rapid 
escalation of high overvoltage. Pflantz (32) found that 
prestrike transients can have crest voltages up to 7.0 pu of 
the system peak, with an oscillating frequency band that 
spans from 60 Hz into the megahertz region. 

The combination of capacitance switching, chopping, 
and prestrike creates many problems in the design of 
high-voltage distribution. If capacitance on the down- 
stream side of an interrupter is small, destructive voltage 
transients can occur from current chopping. When this 
load-side capacitance is large, overvoltages may result 
from capacitance switching or a prestrike event. Conse- 
quently, the placement of capacitance within a mine power 
system is critical, and this will be discussed later in this 
chapter. 

It is interesting to note from the foregoing discussion 
of ac switching transients, that the resulting overvoltage 
in each case can crest at about 7.0 pu of the system peak 
voltage or more. The oscillation frequency provides the key 
to distinguishing among the three types: the system 
natural frequency indicates capacitance switching or 
chopping (generally less than 10 kHz), and much higher 
frequencies indicate prestrike (often these have frequency 
components greater than 100 kHz). 

Direct Current Interruption 

The dc systems are also subject to abnormal voltage 
transients from circuit openings. Unlike ac, where inter- 



C 



Breaker 




Figure 11.15.— Circuit to demonstrate voltage transients in 
dc systems. 



Current through breaker 



System 
voltage 




Fault Contacts 
occurs part 



Voltage across breaker 

Figure 11.16.— Transient overvoltage resulting from current 
interruption on dc system. 



ruption generally occurs at a current zero, dc must be 
forced to zero. Large amounts of magnetic energy can be 
stored in the system inductance, and any sudden current 
decrease can result in an overvoltage. The energy might be 
transferred to capacitance as in ac circuits, but the energy 
can be dissipated in an arc. 3 The arc voltage tends to drive 
current in the direction opposite to the source, forcing the 
circuit current to zero. 

Figure 11.15 shows a simple dc circuit to illustrate the 
overvoltage phenomenon associated with arcing. The over- 
voltage could be caused by a switching operation to clear a 
short circuit. The relationship for voltage and current is 

L^ = V s - v A , (11.12) 



where L = system inductance, H, 
i = system current, A, 
di/dt = rate of current change, A/s, 
V s = source voltage, V, 
and v A = arc voltage, V. 

To stop current flow, the arc voltage must be greater than 
the source potential (fig. 11.16). As the source is constant, 
the peak transient voltage is proportional to the rate of 
current decrease, that is, the faster the decrease, the 
higher the voltage produced. 



3 Personal communication from E. K. Stanek, West Virginia University, 
Aug. 1977. 



287 



The same reasoning can also be applied to interrupt- 
ing dc motor operation as well as other current flows. For 
instance, if a dc motor current is terminated, there exists 
a high rate of current change, which can develop a voltage 
across the system inductance between the source and the 
motor contactors, equal to Ldi/dt. A dramatic example of 
this would be dropping the contactors on a large trolley 
locomotive while drawing full-load current. 

General Switching Transients 

Switching transient problems are not restricted to 
main power components. In either ac or dc systems, small 
tripping and relay coils, when coupled with a very small 
capacitance, can exhibit high voltages even though they 
operate in low-voltage circuits (33). Silicon diodes and 
thyristors can create considerable overvoltage by current 
chopping, sufficient to destroy themselves (19). In fact, 
solid-state conversion equipment is constantly in a tran- 
sient state. 



OTHER TRANSIENT PHENOMENA 

Because of the existing protection methods in mine 
power systems, transients resulting from line-to-ground 
faults or accidental contact between two lines of differing 
voltage are minor, although destruction from localized 
heating may be severe. However, if the protective circuitry 
malfunctions, as might occur in the harsh mining envi- 
ronment, the problem can become critical. Local heating 
at the fault site can cause the conductors to melt. As this 
happens, the potential gradient across the conductors can 
be sufficient to strike an arc. The arc will extinguish and 
reignite, all things remaining equal, at each zero current 
crossing. The random fluctuation of arc impedance, as well 
as phenomena related to arc reignition, results in large 
transient overvoltages (33). 

Resonant conditions can result in transients exceed- 
ing 10 times the nominal line voltage (33). Since the power 
system contains inductance and capacitance that are nor- 
mally much greater than system resistance, resonance can 
occur at the natural frequency of the system. Resonance 
may result from the presence of line-frequency harmonics 
(harmonic overvoltage) or from the frequency components 
of other transients (dynamic overvoltage). Although dy- 
namic overvoltage is probably not a frequent form of 
transient voltage, the number of failures resulting from 
this mode are believed by some engineers to be significant. 
When transformers are involved, transient resonant con- 
ditions creating dynamic overvoltages have traditionally 
been given the term ferroresonance. Considering circuit 
component values, the occurrence of these transients in 
mine power systems should be rare, but a possible problem 
area could be a pole-mounted transformer powered 
through a cable feeder. Here, the series-resonant condition 
could easily be corrected by simply changing the cable 
length. 

Restriking ground faults, which are found primarily 
on ungrounded systems, can cause transients several 
times greater than the nominal line voltages. Some sys- 
tems are operated ungrounded because when a line-to- 
neutral fault occurs, little or no fault current flows and the 
system remains operational. Yet some current will exist 
because of system capacitance. To illustrate the effect of 
this capacitive current during faulting, figure 11.17A 
shows a normal ungrounded system and figure 11.17B 





Figure 11.17.— An ungrounded system, showing capacitive- 
current flow. 




A B 




VCG = ^CA 



Vbg "^ba 



> >Ib+Ic 

Figure 11.18.— An ungrounded system, with fault on phase 



provides the voltage and current phasors. Consider that a 
line-A-to-ground fault occurs as shown in figure 11.18A, 
where I F is a small fault current but sufficient to support 
an arc. The arc may extinguish itself at a current zero, but 
when this happens line-to-line voltages are trapped on 
lines B and C. The system voltage will thus be offset on 
these lines (fig. 11.18B). In the oscillatory return to steady 
state, it is possible to get restrikes of the fault current and 
further self interruption: in other words, a large voltage 
can build up. This is a major reason why systems should 
not be operated ungrounded, unless stern overvoltage 
precautions are heeded. 

Since coal mine power systems are required to use 
resistance grounding for portable and mobile equipment, 
restriking ground-fault transients are also rare, except 
when the protective circuitry malfunctions. A specific case 
would be an open grounding resistor, where uncleared 
faults could cause transients. 



TRAVELING WAVES 

The discussion of transients has considered only cir- 
cuits that have lumped resistance, capacitance, and induc- 
tance, except for circuits in prestrike conditions. In many 
circuits, transient behavior can be accurately predicted 
even though these parameters are distributed (19). But 
there are other power-system portions where circuit- 
element concentration will result in too large an approxi- 
mation. An outstanding example is a transmission or 
distribution line, be it overhead or cable (16). Fortunately, 



288 



these exhibit a certain resistance, inductance, and capac- 
itance per unit length of line, respectively, R, L, and C. For 
analysis, the line can be divided into small but finite 
elements, as shown in figure 11.19 (resistance and leakage 
are neglected) {16). 

The voltages created by transients can have a rise 
time in microseconds; in other words, they rise from zero to 
peak in that time. When very fast voltage changes occur, it 
is often best to analyze the system in terms of traveling 
waves, rather than by conventional methods (16). To show 
why this is so, consider figure 11.19, given that the switch 
has just closed. Conditions existing at the source end are 
not immediately observed at the load end because it takes 
time for the voltage-current conditions to pass through 
each LC segment. The movement of these conditions with 
time is known as traveling waves, and a characteristic 
feature of a circuit with distributed impedance is its 
ability to support these waves of voltage and current (19). 

From analysis of this circuit, where Ax is made very 
short, it can be shown that the voltage-current relation- 
ship for each incremental section is (19) 



dv 



= VF di 



or 



■ w = 



z i, 



where v = voltage existing in incremental element, V, 
i = current existing in that element, A, 
L = total inductance of line, H, 
C = total capacitance of line, F, 

and Z = surge impedance or characteristic imped- 
ance of line, U. 

The propagation velocity, U, of voltage and current is (19) 



U = 



dx 
dt 



1_ 

LC 



(11.14) 



For open lines, the propagation velocity is approximately 
the speed of light, 1,000 ft//xs; with solid-insulation cable, 
the speed is about 500 ft//xs (22). 

To illustrate the time behavior, suppose an overhead 
line with 400-ft surge impedance, as depicted in figure 
11.20A, is hit with an 100-kV step function at time t = 
(16). At this instant, voltage and current exist only at the 
source end and nowhere else on the line (fig. 11.20.B). At 
t = 1.0 /is, the voltage-current conditions have propagated 
down the line for 1,000 ft, and between zero and 1,000 ft, 
the voltage is 100 kV with the current 



v 

z: 



100,000 
400 



= 250 A. 



Beyond 1,000 ft, voltage and current are zero (fig. 11.200. 
At t = 4.0 ^s, the surge has moved 4,000 ft, voltage and 
current being 100 kV and 250 A to the left of that point 
and zero to the right (fig. 11.20D). 

If the line shown in figure 11.20A was of infinite length, 
the bundle of energy would theoretically travel forever. 
Because conductors have finite length, problems occur at the 
line end or at a discontinuity, which is a point where the 
surge impedance changes (16, 19). At either location, the 




AL| Al_2 AL3 



AC, ± AC 2 ;L AC3J; 




AL5 AL 6 



^5± 



ALf LAx 
AC 4 =CAx 

Figure 11.19.— The distributed inductance and capacitance 
of two-wire line shown as incremental sections. 



So 



A 100 kV - 



1,000 2,000 3,000 4,000 5,000 
DISTANCE, ft 



(11.13) C 



D 



m 




W////MWW//, 




.. nn 


W/////////£W^WM 







F 



Figure 11.20.— Demonstration of traveling wave on 
overhead line. 



strict proportionality between the voltage wave and the 
associated current wave must be satisfied by natural adjust- 
ment. Reflected and refracted waves are the result. The 
reflected wave propagates back down the line and is super- 
imposed on the original or incident wave, whereas the 
refracted wave travels beyond the discontinuity. The re- 
flected and refracted amplitudes are such that the voltage- 
to-current proportionalities are preserved, as dictated by 
equation 11.13 and the surge impedances of the lines on 
which they travel (19). Expressed mathematically, 



+ v, = v. 



+ 1, = u 



(11.15a) 



(11.156) 



where 



vi = Z^, 


(11.15c) 


v 2 = -Z^, 


(11.15d) 


v 3 = Z 3 i 3 , 


(11.15e) 



289 



and v^iiiZi = conditions for incident wave, V, A, Q, 
v 2 ,i 2 ,Z 2 = conditions for reflected wave, V, A, Q, 
v 3 ,i 3 ,Z 3 = conditions for refracted wave, V, A, fl. 

The assumption is that energy is conserved. The reason for 
the minus sign indicated in equation 11.15d is that i 2 is 
traveling in a minus-x direction, and thus has a sign 
opposite to v 2 . By combining the above equations, expres- 
sions for the reflected and refracted voltage-wave magni- 
tudes in terms of the incident wave can be obtained: 



-z x 



+ Z, 



2Z„ 



lz 3 + zj v 



(11.16a) 



(11.166) 



Traveling-wave behavior under reflection and refraction 
can be demonstrated using the preceding example and 
terminating the overhead line by either an open circuit, 
short circuit, or a line with a different surge impedance. 
Figure 11.20.E illustrates the conditions if the line end 
is open-circuited. At the instant the incident waves reach 
the end, the current at that point must be zero, as equation 
11.156 relates, 

i, + i 9 = 



or 



lo = -u 



but, by equation 11.15d, 

v 2 = -Zjia = Z^ = v x . 

In other words, the reflected current wave will have a 
magnitude of - i x , and the reflected voltage wave will have 
v x -. Moving to the left, both will superimpose on the 
incident waves. Therefore at t = 6.0 /xs, as shown in the 
figure, the voltage from 4,000 to 5,000 ft is 200 kV and the 
current is zero. 

Instead of an open circuit, consider that the line end is 
short-circuited. When the incident waves reach the short 
circuit, the voltage at that point must be zero, or from 
equation 11.15a, 

Vo = -v,, 



yet, by equation 11.15d, 



2 " Z, " Z x ~ h 



In this case, the reflected waves superimposed on the 
incident waves will result in zero voltage, but at two times 
the incident current magnitude. Figure 11.20F shows the 
conditions at t = 60 ^s, with current to the right of 4,000 
ft as 500 A. 

Now suppose that the overhead line is terminated by 
a cable with a 50-Q surge impedance, a typical value for 
feeder cable. The situations before and after the incident 
waves reach the junction are shown in figure 11.21 (19). 



Zi 



4 



"1 



-k- 



i, 



2, Z 2 

Junction 
Before 



-^— X- 



z, 



Junction 
After 



Figure 11.21.— Incident waves being reflected and refracted 
at discontinuity. 



For the reflected wave (equation 11.16a), 
/50 - 400\ 



450 



100 kV = -78 kV 



and (equation 11.15 a 7 ), 

(-78,000) 



la = 



400 



= 195 A. 



For the refracted wave (equation 11.166), 

/2 (50)\ 



Vo = 



450 



-J 100 kV = 22 kV 



and (equation 11.15e), 



22,000 A - 
x 3 " "SOT- = 445 A " 



Therefore, the surge that penetrates the cable has a 
marked reduction of voltage magnitude. This benefit is 
employed in power systems to protect equipment from 
surges coming down connected lines (19); instances in- 
clude machine trailing cables in surface mines and feeder 
cables in underground mines. It should be noted that for 
this case, the refracted waves propagate at 500 ft//xs, 
whereas the reflected waves are traveling at 1,000 ft//*s. 

The foregoing discussion can be expanded to any line 
termination, including those with more than two connec- 
tions. Within a specific line, reflections can occur at both 
ends, causing the traveling waves to move back and forth. 
In this theoretical model, resistance and leakage have 
been neglected so the energy of the traveling wave is 
maintained. In practical circuits, the presence of resis- 
tance and leakage means losses are incurred when current 
flows (19). These losses serve to attenuate the magnitude 
of the waves as they propagate. Even though the preceding 
demonstrations were simplified through use of a step 
function and by ignoring losses, the concepts can be 
applied to any waveform with a very fast voltage rise time. 

Two important points can be extracted from this brief 
outline of traveling waves (16). First, if a line is open 
ended, any terminal equipment on the line may experi- 
ence a potential up to two times higher than that of the 
traveling wave that produced it. Next, it is a common 



290 



practice to protect personnel working on exposed power- 
lines by placing protective grounds on each power conduc- 
tor. Consider what would happen if a traveling wave 
caused by a lightning stroke was on the line and a person 
was touching the line between the surge source and the 
protective ground. The person would be exposed to the 
incident voltage for the time it took the wave to travel from 
that location to the protective ground and back. It is then 
obvious that protective grounds should always be installed 
right at the worksite, or at least between probable surge 
locations and the workers. 



ELECTROMAGNETIC PHENOMENA 





A 2 charged conductors B 3d conductor presence 

Figure 11.22.— Electric field between conductors. 



It should now be apparent that electrical transients 
are in essence electromagnetic phenomena. According to 
Faraday's law, transients existing on a specific circuit can 
produce electrostatic and electromagnetic induced volt- 
ages on nearby or associated circuits. 

It was shown in chapter 3 that whenever two charged 
conductors are separated, an electric field and a potential 
difference exist between them. The charge is related to the 
potential difference by the proportionality of constant 
capacitance. The physical arrangement between the con- 
ductors is shown in figure 11.22A (19). As shown by figure 
11.22.B, when another conductor is placed in the same 
space, it distorts the original electric field and there will 
be a charge separation on the third conductor surface. The 
conductor will assume a potential somewhere between the 
original two conductors. The relationship will also estab- 
lish capacitances among the three conductors (19). 

Suppose that conductors 1, 2, and 3 are, respectively, 
high-voltage power, grounding, and ground-check conduc- 
tors. If the potential difference between the power and 
grounding conductors suddenly changes, an electromotive 
force is induced in the ground-check conductor, which can 
cause current flow. This redistribution of charge could 
operate a relay, thereby nuisance-tripping a circuit 
breaker. If the third conductor is for some other control, 
monitoring, communication, or grounding circuit, the re- 
sult, beyond false relay operation, could be incorrect ma- 
chine operation, erroneous meter readings, communica- 
tions noise, or injury to personnel. 



TRANSIENT-INDUCED FAILURES 

Deteriorating electrical insulation affects the entire 
mine power system and may jeopardize its safe operation. 
Whether the dielectric is in a transformer winding, motor 
winding, portable cable, or rectifier, it is a critical factor in 
the safe, economical, and reliable operation of any mine 
power system. Disturbances that threaten to compromise 
the integrity of the power system must be eliminated at 
the source. This is why attention must be paid to the 
causes of electrical transients as well as to their elimina- 
tion. Consequently, it may be helpful to examine the effect 
of abnormal voltages on dielectrics. 

Each type of insulation or dielectric is designed for a 
safe maximum applied voltage and a transient overvolt- 
age. The transient overvoltage rating is given in terms of 
BIL, the basic impulse insulation level. The most common 
BIL measurement is the 1.2 x 50 wave test (fig. 11.23), 
where the voltage impressed across the dielectric reaches 
its peak in 1.2 us (22). Thus a dielectric with a 95-kV BIL 
rating can safely withstand a 1.2 x 50 pulse of 95 kV 



100 - 




1.2 



20 30 40 
TIME, fis 



Figure 11.23.— A 1.2 x 50 wave test used for BIL measure- 
ment. 



peak. The peak voltage in the 1.2 x 50 test is considered 
more severe than the transients actually found in mine 
power systems. 

Dielectric deterioration is created in large part by the 
rise time of the transient, as well as the crest magnitude. 
Furthermore, the greater the overvoltage pulse width 
(duration), the greater the probability for failure, since 
more energy is involved. These voltage anomalies break 
molecular bonds in the dielectric, which reduces its effec- 
tiveness. If the overvoltage contains sufficient energy, the 
dielectric can fail immediately; however, this is usually 
not the case. Instead, the insulation is progressively 
weakened until it finally fails, resulting in a line-to- 
ground or line-to-line fault. If the weakened insulation is 
in a portable cable or splice, a considerable personnel 
safety hazard arises since the insulation appears to be 
functional when in reality a lethal potential may exist on 
the cable surface. Although the deteriorating dielectric 
may not present a direct safety hazard, a complete failure 
can, because it may cause an explosion or arcing. In either 
case, the equipment faces serious problems in terms of 
repair, replacement costs, and downtime. 

Winding Response 

The physical construction of equipment may increase 
its susceptibility to transient failure; for instance, motor 
and transformer windings often fail at the end of a coil 
because of increased electrical gradient. Figure 11.24 can 
be used to explain why this can happen (19). It shows the 
distributed nature of the winding inductance and capaci- 
tance, where capacitance is assumed to be uniformly 
divided among the windings and to consist of capacitance 
to ground and to adjacent turns; resistance is neglected. 
The winding neutral may be grounded depending upon the 
position of the switch. 



291 



From an analysis of the circuit, the following general 
equation can be obtained: 



u at 2 " * dx 2 dt 2 



d 4 V 1 d 2 V 



LdX 2 



(11.17) 



where C = winding capacitance to ground, F/m, 
K = winding series capacitance, F/m, 
L = winding inductance, H/m, 

and V = voltage applied across the winding, V. 

By simplification, this can be used to solve for the response 
of the winding to a surge. Consider that surge waveform is 
a step function with amplitude V. If the neutral is 
grounded, at the instant the step function hits the wind- 
ing, the initial voltage distribution across the winding is 



V. = V 



sinh(ax) 
sinh(af) 



(11.18) 



Line end 



'o^wmmmwm^ 



Neutral 



i >J L, lJ L^ U C >J L, J L, >J L> iJ L> .-J U i 



i 




Figure 11.24.— Equivalent circuit of multiturn winding show- 
ing distribution inductance and capacitance. 



and for the ungrounded neutral, 

cosh(ax) 



V = V 



cosh(af) ' 



(11.19) 



where a = 



VI 



I = winding length, m, 
x = distance between neutral and a point on the 
winding, m, 
and V x = voltage at that distance x, m. 

For a specific a, the initial voltage distribution across the 
winding in response to the surge can be plotted as shown 
in figure 11.25 (19). This figure relates that as a increases, 
the distribution becomes nonuniform. When a = 10, 60% 
of the voltage is initially impressed across the first 10% of 
the winding, with 75% across the first 20%. Therefore, 
under surge conditions, very high stresses can occur on the 
first few turns, and if precautions are not taken, trans- 
formers and motors can fail by breakdown of the turn- 
to-turn insulation in this region. 

Coupling Through Transformers 

When a fast rise-time voltage surge hits a trans- 
former, as just shown, the parameter of concern is the 
winding capacitance. Series capacitance and capacitance 
to ground exist not only in the primary, but in the 
secondary winding as well. Figure 11.26A approximates 
this situation for a two-winding transformer, and the 
equivalent circuit shown in figure 11.26B forms a crude 
capacitive voltage divider (19). When a change of voltage is 
applied to the primary, the voltage divides inversely with 
capacitance (15, 35): 



V, = 



Cx 



Ci + C. 



V. 



(11.20) 



where V p = magnitude of transient voltage impressed on 
primary, V, 
V 3 = magnitude of transient voltage transmitted 

to secondary, V, 
Cj = primary-to-secondary winding capacitance, 
F, 
and C 2 = secondary-to-ground winding capacitance, F 



100 



Q 


80 


^ 




D 




O 


60 


LC 




O 




o 


40 


H 






20 



























^(=0 








,3=5 


a=iN 






a=i0 











1 0.8 0.6 0.4 0.2 



Grounded neutral 



too 

80 
60 
40 
20 




a = 



\ 










^ 




a=i 






\ 
















aT^ 




a-\0 




s<£=5 







1 0.8 0.6 0.4 0.2 



Ungrounded neutral 



Figure 11.25.— Initial voltage distribution across uniform 
winding from step function. 



Secondory 

'mm. 




WW/r 




Core- 



B 



Figure 11.26.— Capacitive coupling of transient voltage 
through two-winding transformer. 



Obviously, the voltage transmitted from the primary to 
the secondary is not tied to the transformer turns ratio. 
Typical values are 35% to 40% of the impressed primary 
voltage (15). If the transformer is a step down, for example, 
from distribution to utilization, the result can be low-side 
per-unit voltage levels much greater than those that 
existed on the high side. 



292 



TRANSIENT PROTECTION 

Ideally, the elimination of transient-voltage problems 
begins with an excellent power-system design applying 
time-tested principles. The basic goal is to avoid the 
situations covered in this chapter. However, even this ideal 
situation can address only normal conditions; unpredict- 
able abnormal conditions can still arise, producing de- 
structive transient overvoltages (19). To address this prob- 
lem, additional overvoltage control must be placed in the 
mine power system through use of such protective devices 
as surge arresters, surge capacitors, shielding, and circuit 
arrangements. 

The role of these protective schemes is to ensure that 
equipment dielectric strengths are not exceeded, so that if 
a transient attempts to raise the voltage above an insula- 
tion withstand level, the protective device will exert a 
clamp or restraint to maintain the voltage within accept- 
able limits (19). Effective transient voltage suppression is 
basically the dissipation of transient energy. 

Surge Arresters 

The simplest form of overvoltage device is the spark 
gap, which is essentially two conductors separated by air, 
as in the tips of two rods where one side is connected to a 
powerline, the other to the neutral or earth. The spacing 
between the conductors establishes a specific dielectric 
strength so that voltage above that level will cause the gap 
to spark over. The spark gap has no effect on normal 
system operation, but its main disadvantage is that once 
an arc occurs, a fault is created on the system, which 
remains until the gap is deionized. Often, the attendant 
current flow or follow current can only be interrupted by a 
circuit breaker or a fuse. Surge arresters, formerly called 
lightning arresters, use this spark-gap principle to clip the 
peak of a voltage and divert the excess current to ground. 
They also contain a device to interrupt the follow current 
(31, 34). There are two common surge-arrester types, 
expulsion and valve; they differ in the scheme they use for 
interruption. 

The expulsion surge arrester extinguishes the arc in a 
manner similar to an expulsion fuse (see chapter 9); that 
is, an overvoltage establishes an arc across the spark gap 
and also across a gas-evolving material (usually organic). 
Ignition of the material causes the expulsion of gas, which 
blows out the arc. The operation has three disadvantages. 
First, some gas-producing material is destroyed during 
each operation, and only a limited number of interrup- 
tions are available. Second, because of the gaseous dis- 
charge, care must be taken in placement and installations 
are restricted to outdoors. Lastly, the arrester has an 
assigned current-interrupting rating and cannot be used 
on circuits that have a greater available fault current than 
this rating. 

In valve surge arresters, the spark gap is in series 
with a nonlinear resistor or valve block, as shown in figure 
11.27. A property of the block, which is commonly made of 
silicon carbide, is that the resistance diminishes sharply 
as the voltage across it increases. Tb increase gap effi- 
ciency, a number of short gaps are used because these 
spark over more consistently and in less time than one 
long gap. On an overvoltage, the gaps spark over and the 
valve block operates in its highest conductance to pass the 
surge current safely to ground. After the surge voltage 
diminishes, the block changes to a low-conductance mode 



Line 




Spark gap 



Valve block 



— Ground 
Figure 11.27.— Basic valve surge arrester. 



to limit the follow current, such that the gaps can provide 
an interruption. The valve surge arrester has none of the 
disadvantages of expulsion types, and it is used exten- 
sively for equipment protection, especially on distribution 
systems. The balance of this section will thus cover only 
the valve units. 

Four important parameters are connected with the 
proper application of surge arresters (31 , 34): 

1. Voltage Rating. The power frequency sparkover 
voltage is the lowest rms 60-Hz ac voltage across the 
arrester at which it will perform the operating cycle. This 
level is 1.5 times the arrester voltage rating for arresters 
rated at 60 kV and below. 

2. Sparkover Voltage. The highest crest voltage at 
which arcs will form across the spark-gap electrodes, 
initiating the operating cycle. 

3. Discharge Current. The current through the arrester 
created by the overvoltage immediately after sparkover. 

4. IR Discharge Voltage. The voltage formed across the 
arrester during the discharge of surge current. 

Ideally, gap sparkover should occur on any dangerous 
system overvoltage and ignore all minor and harmless 
transients. Proper sparkover requires high-speed response 
to fast rise-time wave fronts (as in lightning surges) and 
consistent response to slower rates of voltage rise (as in 
system-generated surges). Both requirements are satisfied 
by electrical grading of the spark-gap structure, which 
consists of shunting each gap with a high resistance. 
Figure 11.28 shows the technique in simplified form (31). 
After sparkover, the IR discharge voltage occurs, being 
equal to the product of the discharge current (I) and the 
arrester discharge-path resistance (R). As discharge cur- 
rent may be very large, the discharge voltage may equal or 
exceed the sparkover voltage. Thus, protected equipment 
is exposed to both the sparkover and IR discharge voltages, 
and the system insulation withstand ability must be 
safely above both. 

To establish an IR discharge voltage, it is important to 
recognize the magnitude of possible discharge currents. A 
surge arrester is likely to be exposed to a wide discharge 
range, but experience from field measurements has shown 
that discharge currents typically range from 1,000 to 
2,000 A, that 5.0% exceed 9,000 A, and that only 1.0% 
exceed 20,000 A (31). Even though 20,000 A is rare, it is 
often used as a worst case to estimate the discharge 
voltage. 



293 



Gap 
electrodes 



Gap 
electrodes 



Nonlinear 

grading 

resistor 

Nonlinear 

grading 

resistor 




Gap 
electrodes 



Nonlinear 

grading 

resistor 



Valve block 



Figure 11.28.— Surge arrester with nonlinear resistance 
grading to equalize each gap structure. 



There are three classes of valve-type arresters avail- 
able (34): 

1. Distribution-class arresters have the lowest cost and 
are satisfactory for general equipment protection pur- 
poses. Typical voltage ratings range from 1,000 V to 18 kV, 
and these arresters can withstand a 65-kA current surge. 
(Surge arresters are generally used on high-voltage sys- 
tems in mining, but these devices are available with 
ratings as low as 50 V.) When protecting rotating machin- 
ery and dry-type transformers, the arresters must be the 
low-sparkover rotating-machinery (RM) type because of 
low BTL's. 

2. Intermediate-class arresters have lower sparkover 
and IR discharge voltages than do distribution-class ar- 
resters. Available voltage ratings are from 3,000 V to 73 
kV, and the arresters can typically withstand a 65-kA 
surge current. The cost is about five times that for a 
distribution arrester. 

3. Station-class arresters can handle discharge cur- 
rents up to 100,000 A and are available for almost any 
distribution or transmission voltage application. They are 
considered to provide the best possible protection, but cost 
about twice as much as intermediate-class arresters. Both 
station-class and intermediate-class arresters have 
pressure-relief systems; if stressed by a surge beyond their 
capability, the internal pressure is vented to prevent 
housing rupture. 

Surge Arrester Applications 

The factors to consider when selecting an arrester 
class are the degree of transient exposure and the impor- 
tance of the equipment being protected. This is basically 
an economic question, but in general, intermediate or 
station arresters are justified for surface substations, with 
distribution-class arresters being suitable for distribution 
and utilization equipment protection. 



An arrester voltage rating of a certain class has 
associated sparkover and IR discharge voltages. System 
voltage, as well as the method of system grounding, affects 
the voltage that the arrester is exposed to and therefore 
the selection of the arrester voltage rating. Consequently, 
once an arrester voltage rating is set, the system insula- 
tion withstand ability must be coordinated with it. This is 
most often related as a BIL for the equipment being 
protected. 

The voltage-rating selection is affected by the system 
grounding categories: effectively grounded or noneffectively 
grounded. The coefficient of grounding can be employed to 
find which category is being used. The coefficient of 
grounding is defined as the percent ratio of the highest 
rms line-to-ground voltage existing during a line- 
to-ground fault to the nominal line-to-line voltage (31). If 
the ratio does not exceed 80%, the system is termed 
effectively grounded. Solidly grounded and typical low- 
resistance-grounded systems are in this class (the neutral 
potential remains rather constant during line-to-neutral 
faults). However for high-resistance and ungrounded sys- 
tems, the occurrence of a line-to-neutral fault can shift the 
neutral to near the faulted line, with the potential to 
ground of the other two lines approaching line-to-line 
system voltage. The coefficient of grounding here can be 
from 80% to 100%. These systems are termed noneffec- 
tively grounded. Resistance-grounded mine power systems 
for portable and mobile equipment are included in this 
category. 

In either grounding case, the arrester voltage rating 
should be above the possible exposed crest voltage; if not, 
a disruptive discharge might occur. 

Therefore, on effectively grounded systems, the ar- 
rester is sized to maximum expected line-to-ground volt- 
age, whereas maximum line-to-line voltage is used for 
noneffectively grounded systems (2). To allow for the 
expected increase due to voltage-regulation compensation, 
the arrester voltage rating should be 5% to 10% above 
these values. Tables 11.1 and 11.2 list the recommended 
sizing for resistance-grounded mine power systems (3, 34). 
The first table refers to station and intermediate arresters, 
the second to the low-sparkover distribution class. Note 
that the transformer BIL's specified are according to ANSI 
C57. 12. 00-1973 (1) and may be too low for some mining 
applications (see chapters 12 and 13). 

The equipment protection from voltage surges can be 
verified for any arrester selected. Full coordination re- 
quires checking the arrester performance over a full time 
range for an expected surge (1). However the following 
quick guidelines will ensure safe protection (34). 

• The insulation BIL ratings of equipment should be 
at least 20% greater than the arrester sparkover voltage. 

• The BIL rating should be above the IR discharge 
voltage of the arrester. 

As mentioned earlier, a discharge current of 20,000 A to 
establish IR discharge voltage may be used as an approx- 
imation of worst case conditions. Tables 11.1 and 11.2 also 
provide typical sparkover and IR discharge voltages for 
arresters used in mining service (34). However, manufac- 
turer catalog values should be consulted for actual appli- 
cations, as differences exist among products. Additional 
details of surge-arrester equipment protection are pro- 
vided in chapters 12 and 13. 

Another important point about maximum exposed 
surge voltage concerns the arrester connections to line and 



294 



Table 11.1.— Recommended station and intermediate surge arresters for resistance-grounded mine power systems to protect 

oil-immersed transformers 



Insulation 

class. 

kV 



System 

voltage, 1 

V 



Transformer 
BIL, 2 kV 



Arrester 
rating, 



Front-of-wave 

sparkover, 3 

kV 



IR discharge 

voltage 3 
(20,000 A) kV 



Maximum 

3-phase 
line-to-line 



Nominal 


Maximum 


Power 


Distribution 


l\V 


SC 


IC 


SC 


IC 


voltage, ' 


2,400 


2,540 


60 


45 


3 


13 


12 


7.8 


12 


3,000 


4,160 


4,400 


75 


60 


6 


19 


21 


15.5 


20 


6,000 


4,800 


5,080 


75 


60 


6 


19 


21 


15.5 


20 


6,000 


7,200 


7,620 


95 


75 


9 


30 


31 


23 


27 


9,000 


12,470 


13,200 


110 


95 


15 


51 


51 


39 


45 


15,000 


13,200 


13,970 


110 


95 


15 


51 


51 


39 


45 


15,000 


13,800 


14,520 


110 


95 


15 


51 


51 


39 


45 


15,000 


14,400 


15,240 


110 


95 


18 


NA 


61 


NA 


54 


18,000 


14,400 


15,240 


110 


95 


21 


70 


NA 


54 


NA 


21 ,000 



2.4. 
5.0 

8.7. 
15., 



NA Not available. 

1 Maximum system voltages are from ANSI C84.1-1970 (4). 

2 BIL ratings are from ANSI C57. 12.00-1 973 for oil-immersed transformers (1). 

3 SC, station class; IC, intermediate class. 



Table 1 1 .2.— Recommended distribution-class, RM-type, surge arresters for resistance-grounded mine power systems to protect 

rotating machinery and dry-type transformers 



Insulation 

class, 

kV 



System 
voltage, 1 V 



Nominal 



Maximum 



Transformer 

BIL, 2 

kV 



Arrester 

rating, 

kV 



Front-of-wave 

sparkover, 

kV 



IR discharge 

voltage 

(5,000 A), 

kV 



Maximum 

3-phase 

line-to-line 

voltage, V 



2.4. 
5.0. 

8.7. 
15.. 



2,400 

4,160 

4,800 

7,200 

12,470 

13,200 

13,800 

14,400 



2,540 

4,400 

5,080 

7,620 

13,200 

13,970 

14,520 

15,240 



20 
25 
25 
35 
50 
50 
50 
50 



3 
4.5 

6 

3 7.5 
15 
15 
15 
3 15 



13 
17 
22 
24 
44 
44 
44 
44 



10 
15 
20 
25 
50 
50 
50 
50 



3,000 

4,500 

6,000 

7,500 

15,000 

15,000 

15,000 

15,000 



1 Maximum system voltages are from ANSI C84.1-1970 (4). 

2 BIL ratings are from ANSI C57. 12.00-1 973 for oil-immersed transformers (7) 

3 Arresters may occasionally be exposed to voltage above their rating. 



to ground. Conductors extend from each ungrounded 
power conductor to an arrester and from the arrester to 
ground so that in resistance-grounded three-phase sys- 
tems, a minimum of three arresters is needed. For station- 
ary equipment on the surface, an arrester ground bed, 
such as a substation system ground bed, serves as the 
grounding medium (it must be low resistance); on portable 
equipment, the frame would be the ground. As the connec- 
tions exhibit inductance, they should be of No. 6 AWG 
solid-copper conductor or larger and as short as possible 
because the inductance of too long a conductor can render 
an arrester ineffective (19). Sharp bends should also be 
avoided, since a bend substantially increases the effective 
inductance. The inductance of arrester connections that 
adhere to these requirements is estimated at 0.4 /xH/ft, 
with the voltage drop produced during a surge estimated 
at 1.6 kV/ft, using a current wave front of 4,000 AJ/is (31). 
This voltage must be added to sparkover and IR discharge 
voltages to assess protective margins. 

As a general rule, all arresters should be located as 
close as possible to the equipment they are to protect. 
Ideally, they should be across the protected equipment 
terminals, the connections for three-phase systems being a 
wye configuration with the common arrester connection 
grounded. First of all, this location minimizes the possi- 
bility of a destructive surge entering the circuit between 
the protecting and protected devices (19). Second, close 
proximity also reduces the change of surge-voltage ampli- 
fication through refraction and reflection of a traveling 



wave. For instance, consider figure 11.29, where a tran- 
sient voltage surge is traveling along a power conductor 
toward surge-arrester-protected equipment (22). The ar- 
rester is located a distance, d, from the line end. As the 
wave passes the arrester location, the arrester sparks over 
at its protective level, but lets a traveling wave with a crest 
equal to its sparkover value through. The voltage wave 
reaching the equipment terminals is amplified by the 
addition of the incident and reflected wave, with the 
resultant magnitude depending upon the line-end surge 
impedance. Consequently, surge-arrester locations other 
than directly across the equipment terminals can lead to 
higher surge voltages at the protected apparatus than the 
arrester sparkover voltage. The terminal voltage rise will 
be aggravated by a greater separation distance, d. Note 



30 kV 



Surge arrester 




Line termination 



Figure 11.29.— Surge approaching surge-arrester-protected 
equipment. 



295 



that with a wavefront rise time no greater than 0.5 fis, a 
maximum distance of 25 ft is perhaps allowable, but 
shorter distances always afford greater protection (28). 

Capacitors and System Capacitance 

Surge capacitors, also termed RM capacitors, are 
special units with low internal inductance that are used 
extensively to protect rotating machinery and dry- 
insulated transformers (28). This equipment is very sus- 
ceptible to line-end turn-to-turn failures caused by fast 
rise-time wavefronts, and the faster the rise time, the 
greater the probability for damage. Connected across 
equipment terminals in grounded wye, as shown in figure 
11.30, the capacitors serve to limit the rate of rise of the 
transient voltage. Simply, the capacitor has to be charged 
before the overvoltage can be impressed on the system 
dielectric. Transient rise time is then largely determined 
by the charging rate. Coupled with the system inductance, 
the limiting criterion is that at least 10 (is is needed before 
the crest value of the protected-equipment nameplate 
voltage is reached (22). Low internal inductance of the 
capacitor is important because the presence of series 
inductance in the capacitor circuit deteriorates the wave- 
sloping action. Accordingly, the capacitors shown in table 
11.3 have been standardized for this kind of protection 
(22). In combination with the recommended low-sparkover 
distribution-class surge arresters (table 11.2), the crest 
voltage of transients is considered restricted to harmless 
values for the utilization equipment (31). 

Table 1 1 .3.— Commonly used surge capacitors for limiting 

voltage rate of rise on rotating machinery and dry-insulated 

transformers 

Rated equipment Capacitance, 

voltage, V nF 

650 or less 1.0 

2,400 to 6,900 .5 

11,500 or higher .25 



Surge-Impedance Reduction 

Surge capacitors can also be used to control transient 
overvoltages by reducing the system's characteristic im- 
pedance. An increase of system capacitance, as exhibited 
by equation 11.8, can lower the surge impedance and 
therefore the possible peak transient voltage resulting 
from current chopping. A fine line exists here, because too 
much capacitance on the load side of switching apparatus 
can cause capacitive-switching or prestrike transients. For 
instance, prestrike events are dependent upon the capaci- 
tive inrush current during contact closure, and the mag- 
nitude of inrush current is controlled by the amount of 
load-side capacitance (30). With capacitive switching, the 
transient overvoltage is created by the ability of the 
load-side capacitance to hold a charge, thus causing a 
recovery voltage across the breaker so restriking occurs. 
System capacitance is, therefore, a critical factor in tran- 
sient protection. 

Hence, the problem of surge-impedance reduction 
mainly concerns how much capacitance is necessary to 
limit safely the transient overvoltages caused by current 
chopping. Actually, any magnitude below the minimum 
insulation withstand level or, if used, the surge arrester 
sparkover can be considered safe. Yet perhaps the most 
conservative approach would be to limit any chopping 



^nmnmnnn^n 



Series inductance 



Standard 
arrester 



Low 
sparkover 
arrester 




Machine 
frame 



Figure 11.30.— Typical surge protection of rotating 
machinery and dry-insulated transformers. 



event to two times the peak system voltage. The most 
critical portion of the mine power system would be where 
capacitance is minimum, such as the case of a switchhouse 
connected to a power center, as illustrated in figure 11.14, 
or rotating machinery. It can be inferred that current 
chopping is the result of VCB operation, but other 
switching-apparatus types, including current-limiting 
fuses, could be chopping sources. 

Equation 11.8 can be used to select a value of capac- 
itance that will include the capacitance inherent in the 
distribution system, that is, the equivalent line-to-neutral 
capacitance of all devices on the load side of the switching 
apparatus. Considering a three-phase transformer as the 
load, the procedure can be as follows: 4 

1. Determine the allowable peak voltage, V ; 

2. Find the exciting current and assume interruption 
is at peak level, I m ; 

3. Determine the transformer exciting inductance, 

L m ; 

4. Assume all transient energy is absorbed by the 
capacitance, C; then 

5. The necessary per-phase capacitance, in farads, 
referred to the transformer primary circuit is 



C = 



V 2 



(11.21) 



However, a more useful form of this equation can be found 
if (24) 

1. The allowable peak voltage is limited to two times 
the peak system voltage; 

2. The inductance of the transformer and the power 
system is no greater than 20%; and 

3. The peak exciting current is expressed in terms of 
the rated capacity and rms voltage of the transformer. 



With these parameters, 



C = 



10 S (60) 
V 2 f 



(11.22) 



where C 

S 

V„ 



and 



f = 



= necessary capacitance, (i¥, 
= per-phase transformer capacity, VA 
= line-to-line voltage rating of transformer pri- 
mary, V (if it is desired to connect the capac- 
itance across the secondary, secondary line- 
to-line voltage is used), 
power-system frequency, Hz. 



4 Personal communication from E. K. Stanek, West Virginia University, 
Aug. 1977. 



296 



The value resulting from equation 11.21 or 11.22 is total 
system capacitance per phase. If the level is above that 
supplied by the system, additional capacitance might need 
to be added. 

For three-phase systems, the common surge capacitor 
connection is shown here in a wye configuration with the 
center connected to ground. The typical location is directly 
across the protected equipment terminals, as in wave- 
sloping applications (fig. 11.30). Another popular location 
has been at the switching-apparatus load terminals. The 
philosophy here is that the interrupter sees the increased 
capacitance directly, and thus, chopping transients are 
limited more effectively; in other words, transients are 
best eliminated at their source. There is a general feeling 
that in this way the entire system downstream from the 
capacitor location would receive protection. However, 
surge capacitors have an extremely low internal series 
impedance; therefore, at the interrupter load terminals, 
they can be a significant source of capacitive inrush 
current as well as having the ability to exchange transient 
energy effectively to and from the system. Consequently, 
the best location for applying surge capacitors is at the 
protected-equipment terminals. 

Two advantages are gained through the ground con- 
nection: Transient energy is shunted to ground, and com- 
pared with delta connections, a lesser voltage rating is 
necessary. The capacitor working-voltage rating (WVDC) 
should be at least three times the exposed rms system 
voltage, usually line-to-neutral for wye grounded and 
line-to-line for delta connections. 5 However, for the same 
reasons given for surge arresters, it is perhaps best to rate 
surge capacitors in resistance-grounded systems to three 
times the line-to-line voltage. 

Buss (11) and Morley (30) have performed extensive 
tests on actual underground coal mining equipment to 
determine the severity of transients existing on these 
distribution systems. For the most part, these tests in- 
volved recording staged transients on unloaded and loaded 
systems by chopping (tripping the interrupter) and pre- 
strikes (engaging the interrupter). In every instance, the 
system segment was similar to figure 11.31 and consisted 
of a VCB-equipped switchhouse with various lengths of 
cable supplying a power center. Various switchhouses and 
power centers were used, all typical of actual mine instal- 
lations. Beyond the principal goal of uncovering the na- 
ture of any transients, the overall objective was to see if 
surge capacitors were necessary to limit chopping voltage 
transients and what effect they have on prestrike events. 
Buss (11) backed up the actual equipment testing with 
computer simulations but primarily addressed chopping 
events. Morley (30) extended the research to cover pre- 
strike activity. 

In terms of transient protection by surge capacitors, 
the results of both research programs were practically 
identical and apply directly to underground mine power 
systems or similar surface arrangements. Specifically, it 
was found that both grounded wye or ungrounded wye 
surge capacitors do significantly reduce chopping tran- 
sients when applied as given in table 11.3. With capacitors 
removed and practical cable lengths held to a minimum, 
the chopping transient voltages reached substantial levels 
(about 4.5 pu system nominal voltage for a common event) 
but were still below the sparkover level of correctly applied 
low-sparkover distribution-class surge arresters. Thus 



Mine 

distribution 

system 



u 



r^> 



Measurement 
point 



Feeder 

cable 

« «- 



SWITCHHOUSE 



Feedthrough 



Measurement 
point 

4 



rni — uzt 

iH 



Wye 
grounded 



5 Personal communication from E. K. Stanek, West Virginia University, 
Aug. 1977. 



LOAD CENTER 

KEY 
C Surge capacitor -usually 0.25,6<F 
SA Low-sparkover distribution- class surge arrester 

* Most load-center transformations were delta-wye: 
some were wye-delta or delta-delta. Common 
capacity was 750 kVA; other capacities included 
650 and 1,000 kVA. 

Figure 11.31.— Simplified sketch of mine power-system seg- 
ment. 



they remained within the BIL withstand of system equip- 
ment. However, with surge capacitors installed and a 
1,000-ft cable length, some prestrike transient activity 
was present, with peak voltages similar to the maximum 
observed for chopping but again not exceeding the surge 
arrester sparkover. The worst case prestrike activity was 
associated with surge capacitors located at switchhouses. 

The conclusion was that surge arresters alone were 
sufficient transient protection for underground coal mine 
distribution systems. Surge capacitors were unnecessary 
redundant protection, and their general use could result in 
more severe transients than those they are installed to 
correct. Furthermore, as long as surge arresters protect 
each distribution load, VCB's can be employed on any 
mine distribution system without a need to add surge 
capacitors (11). 

However, when surge capacitors are not used, a sig- 
nificant length of cable should be installed between 
switchhouses and loads to limit chopping events to safe 
levels. For general mining applications, 500 ft of SHD 
cable was recommended, although 100 ft would probably 
be satisfactory (11 , 30). When the capacitance is calculated 
from equation 11.21 and compared with cable values (in 
tables 11.4 and 11.5), comparable results are supplied. 
Beyond overall surge-impedance reduction (if lumped cir- 
cuit elements are considered), the normally lower charac- 
teristic impedance of the cable compared with that of the 
switchhouse circuits reduces the crest of voltage traveling 
waves through refraction. 

The above findings were for common system arrange- 
ments in mining. Some installations could still benefit 
from the wave-sloping action of surge capacitors, as for 
example a surface mine where there is a high incidence of 
lightning and the distribution loads are rotating machines 
or dry-insulated transformers. 

Capacitive Charging Current 

Wye-grounded surge capacitor combinations have 
been recommended for several years to prevent severe 
chopping transients. Therefore, mines often employ them 



297 



Table 1 1 .4.— Typical capacitances, in microfarads, per phase 

of power-system components, for shielded power cable SHD, 

SHD-GC, andSHD + GC 

















o 


Insulation class 


kV.. 


5 


8 


15 


23 




AWG: 














a 


8 






0.0607 


— 


— 


— 


< A. 


6 

4 






.0709 
.0836 


— 


— 


— 


h- o> 

9 ° 


2 






.0993 


0.0850 


0.0401 


— 


g * 


1 






.1096 


.0938 


.0456 


0.0345 


< <u 


1/0 






.1202 


.1029 


.0524 


.0388 




2/0 






.1321 


.1131 


.0507 


.0441 




3/0 






.1452 


.1243 


.0661 


.0504 


4/0 

MCM: 






.1600 


.1369 


.0715 


.0546 


< 

X 

<_> 


250 






.1598 


.1368 


.0776 


0588 


350 






.1846 


.1580 


.0844 


.0634 




500 






.2150 


.1840 


.0920 


.0685 




750 






.2410 


.2062 


.0981 


.0743 




1,000.. 






.2740 


.2345 


.1118 


.0789 












NOTE.- 


— Dashes indicate cable 


is not made 











0.040 



.032 



.024 



.016 



.008 































■P 


















i^g 
































































6C 


>°- 


















"ZL§2°— — 
























<2z: 





































1,200 
1,800 



Table 11.5.— Typical capacitances per phase of power-system 
components 



100 200 300 400 500 

MACHINE SIZE, hp 

Figure 11.32.— Capacitance for 2,300- V induction motors; for 
motors up to 7,200 V, value will not vary more than ± 15% of 
above. 



Power-system component Capacitance, /iF 

Nonshielded cable 0.02 to 0.05 per 1,000 ft; typically 

0.03 per 1,000 ft. 

Nonshielded cable in conduit 0.02 to 0.06 per 1,000 ft; typically 

0.04 per 1 ,000 ft. 

Overhead open-wire lines Negligible for line lengths used in 

typical mine distribution. 
Surge capacitors, by insulation 
class: 

600 V 1.0. 

2,000 V 1.0. 

5 kV 0.5. 

8 kV 0.25. 

15 kV 0.25. 

23 kV 0.25. 

Synchronous and induction motors, 
by insulation class: 

600 V 0.032. 

2,000 V to 23 kV See figures 1 1 .32 and 1 1 .33. 



at the high side of every distribution transformer. Such 
extensive use can pose problems in the high-resistance 
grounding system of a mine. 

As stated in chapter 7, the ground-fault current that 
is limited by the grounding resistor should not be less than 
the system capacitive charging current. This requirement 
restricts the amount of capacitance that can be placed on 
a system of a specific size. In other words, the size of the 
mine distribution system is limited by the total connected 
capacitance (11). Some relief is gained through power- 
transformer configurations, because the zero-sequence sys- 
tem is isolated in each voltage level. Nevertheless if there 
is excessive capacitance in distribution, for example, the 
capacitance can discharge during a line-to-neutral fault 
and feed the fault with capacitive ground current in excess 
of current limit. 

Tables 11.4 and 11.5 and figures 11.32 and 11.33 
provide typical system capacitances to assist in estimating 
a per-phase system capacitance (38). The following can be 
used to compute system charging current per phase: 



o 
u 



o 



U Ql 

O u_ 

o 
rr 
< 

x 
o 



0.020 



.016 



.012 



.008 



.004 












































































































p, 




jb0> 


q aH 


















9U 


2800.. 




















^2g 



































100 200 300 400 

MACHINE SIZE, hp 



500 



I =^ 

co X„ 



(11.23) 



Figure 11.33.— Capacitance for 2,300- V synchronous 
motors; for motors up to 7,200 V, value will not vary more than 
±15% of above. 



where I co = per-phase system charging current, A, 
V ln = line-to-neutral system voltage, V, 
X co = per-phase capacitive reactance, Q, = l/27rfC , 

and C = lumped charging capacitance per phase, F. 

This current can be compared with ground-current limits 
to assess the effect of adding surge capacitors. The values 
in the table can also be compared with the results from 
equations 11.21 or 11.22 to obtain greater understanding 
of chopping events. 



298 



Other Suppression Devices 

The discussion thus far has been mainly concerned 
with surge suppression on high-voltage distribution. 
Transient-related failures can also be severe on mine 
systems below 1,000 V, an outstanding example being the 
destruction of solid-state elements in such equipment as 
ground-check monitors, communications apparatus, and 
power supplies. These transistors, integrated circuits, and 
thyristors are the devices most sensitive to transient 
overvoltages, and problems can occur even through induc- 
tion from transients occurring on power conductors to a 
neutral or communications circuit. Two suppressor types 
already presented also offer effective protection for low- 
voltage systems: valve-type surge arresters and surge 
capacitors (also termed snubbers). In this section, several 
other common protection devices designed principally for 
but not restricted to low-voltage applications will be 
discussed. These can generally be divided into two classes: 
transient suppressors and circuit-shorting devices. 

Transient Suppressors 

Transient suppressors, also called constant-voltage 
devices, are basically nonlinear resistances placed across 
the circuit to be protected (13, 18). They act directly to 
clamp or limit voltage rise much like the valve block of a 
surge arrester, but no series spark gap is used. Typical 
transient suppressors are power zeners and metal oxide 
varisters. 

Powers zeners are primarily for dc protection, working 
on the zener-regulator principle covered in chapter 5. They 
can be used in ac circuits when two devices are placed back 
to back or anode to anode to give bidirectional operations. 
Zeners have the capability to clamp transient voltage rise 
to a level largely independent of the impedance or voltage- 
current characteristic of the transient. The response is 
extremely fast, and the clamping action is very firm. 
Although they are effective for low-energy transients, 
many events common in industrial power systems can 
readily destroy all but the most expensive high-energy 
zeners, which are called avalanche diodes (18). 

Metal oxide varisters (MOV's) are ceramic suppressors 
that use zinc-oxide-based materials such as a zinc oxide 
and bismuth oxide ceramic body (18). Their construction 
provides a voltage-dependent, very nonlinear resistance 
with symmetrical conducting properties. The bidirectional 
breakdown allows their use on either ac or dc circuits. The 
response is similar to that of back-to-back zeners, but the 
clamping action is softer than with zeners, yet faster than 
with valve-type surge arresters (13). As with the silicon 
carbide valve block in surge arresters, transient energy is 
dissipated throughout the entire volume of material, mak- 
ing the MOV a very rugged suppressor. 

MOV's can be used for a wide range of applications from 
power-supply voltage regulation to power-system transient 
suppression, including high-voltage systems. Rated voltages 
extend from 22 V to the thousands (18). A very popular 
application in mining is to suppress transients across thyris- 
tors in solid-state motor starters, where an MOV is mounted 
across each thyristor. A projected employment at this time is 
for direct replacement of valve surge arresters in ac distri- 
bution and dc trolley-line applications. 



Circuit-Shorting Devices 

A circuit-shorting device or "crowbar" can be de- 
scribed as a device such as a spark gap, gas-discharge tube 
or thyristor that senses a high voltage and throws a short 
circuit or low resistance across the line (18). The low 
resistance is not removed until the current through the 
crowbar is brought to a low level; hence, these devices need 
to have the power removed momentarily before resetting 
can occur (13). To facilitate this requirement, crowbars are 
often used in conjunction with circuit breakers, where the 
device could be connected in series with a shunt-tripping 
element, and the combination located across the line on 
the load side of the breaker. Sensing an overvoltage, the 
current through the crowbar trips the circuit breaker; 
normal circuit operation is not resumed automatically. 
These quick-acting devices are available for ac or dc 
systems, typically 250 V and below. 

Faraday Shields 

Faraday shields are used to protect the low side of a 
transformer against surge voltages. The shield is a turn of 
nonmagnetic metal sheeting placed between the primary 
and secondary windings; it is insulated from all windings 
and connected solidly to ground (15). This location effec- 
tively destroys interwinding capacitance and substan- 
tially reduces the transfer of surge conditions. The shield 
has the further advantage of practically eliminating inter- 
winding faults. Other advantages of these shields are 
covered in chapters 12 and 14. 

Circuit Arrangements 

As induced voltages on low-voltage circuits from tran- 
sients on high-voltage or other power-system conductors 
can be a serious problem, there are four recommendations 
to reduce the possibility of induction. 6 

1. Separate low- voltage circuits from high-voltage 
systems by a large distance. 

2. Use shielded conductors on the low- voltage circuit 
or maintain shielding between circuits serving different 
purposes. 

3. Twist low- voltage conductors (this effectively can- 
cels many induced voltages). 

4. Place conductors on the high-voltage circuit close 
together. 

Protection of Overhead Lines 

Exposed overhead lines are very susceptible to direct 
contact from a lightning stroke, which obviously produces 
severe transient overvoltages. An excellent means of pro- 
tecting distribution lines from such occurrences involves 
the use of overhead ground wires or static wires (25, 37). 
One or two ground wires are strategically situated above 
and between the power conductors to provide a shielding 
effect, as illustrated in figure 11.34 (27). Here, line a 
bisects a line drawn from the ground wire to the outer 



6 Personal communication from E. K. Stanek, West Virginia University, 
Aug. 1977. 



299 



power conductor; lightning strokes in area 1 are inter- 
cepted by the static wire. Line b is equidistant between the 
outer power conductor and the earth's surface, so that 
lightning streamers in area 2 will discharge to the earth. 
The position of line b in figure 11.34 varies with the 
relative height of the supporting structure, and line c is 
described by an arc whose radius depends upon the size 
and construction of the supporting structure. A stroke in 
area 4 will be borne by the supporting structure if it is 
metallic. Area 3 is the danger region (and may include 
area 4) where lightning flashes will strike the power 
conductor (27). 

Extensive tests and observations have been conducted 
to determine the optimum angle 9, which is measured 
from the vertical up to the line joining the power conductor 
and the static wire. The angle is dependent upon the 
height of the supporting structure, as shown in table 11.6 
(27), but field results have shown that a good average for 
this angle is 30° (37). (Actually, 45° provides satisfactory 
performance if the line is situated on a level surface, but if 
on a hillside, the angle should be decreased from 45° by 
the slope angle of the hill.) Importantly, the design of 
static protection is practically independent of the system 
voltage. Figure 11.35 illustrates two approaches for shield- 
ing overhead lines by static wires when the support 
structure is wooden (37). 



Table 1 1 .6.— Protective angle versus structure height 

Tower height, ft Angle (Q), deg 

50 45 

100 25-30 

150 10-12 



In addition to intercepting direct strokes, the static 
wires perform another function: when static wires are 
struck by lightning, the resulting surge current is imme- 
diately halved since the impulse travels in both directions 
from the point of contact. The magnitude of the induced 
voltage on the static wire is thereby halved also, leading to 
a similar reduction in the induced potentials seen by the 
power conductors. 

Adequate grounding must be used in conjunction with 
the overhead static wires. The system generally consists of 
downleads extending from the static wire to grounding 
electrodes so that induced voltages are minimized by 
conducting stroke currents to earth as quickly as possible. 
This also helps to reduce the surge impedance of the 
overhead static-wire system. A grounding electrode may 
be provided by driving several ground rods at the base of 
each supporting structure and connecting them to the 
static wires via the downlead (23). An alternative method 
is based upon the use of a counterpoise, consisting of one or 
more buried horizontal conductors located at the base of 
each tower (5). The counterpoise may be a single continu- 
ous conductor buried directly beneath the power conduc- 
tors and running from tower to tower, or it may consist of 
several short conductors radiating outward from each 
tower base (9). The most common technique in mines is to 
wrap heavy wire around the base of each wooden pole 
before it is placed in the hole, the so-called butt wrap. 
Whichever technique is selected, the grounding electrode 
must always have as low a surge impedance as practical, 
and the downleads must adhere to the guidelines stated 
earlier for surge arrester connections. As a general rule- 
of-thumb, downleads and grounding electrodes should not 



KEY 
• Static wire 
o Line conductor 
9 Protective angle 




'////// 



v//////, 



'///////, 



v////////. 



•'//////A 



Figure 11.34.— Overhead ground-wire shielding for low and 
high distribution towers. 



Wood braces. 
If steel braces, 
make allowances 
for loss of 
clearance. 



, Static wires -. 
Possible cross 




Insulator 
spacing 



Insulator 
spacing 



H-frame structure 



Single pole 



Figure 11.35.— Static-wire-protection designs of wooden 
support structures using 30° protective angle. 



be spaced apart more than 500 ft. Considering normal 
spans in mining, this translates to grounding the static 
wire no more than every other supporting structure. 

A few additional comments are needed on the appli- 
cation of static wires in mining. At the voltage levels used 
in conjunction with overhead lines to distribute mine 
power, the normal separation distance used for insulation 
between power conductors and static wires is perhaps 
insufficient to prevent a lightning stroke that hits a static 
wire from jumping directly to the power conductors (37). 
Often, the static wire is connected to earth only at the 
system ground bed at the substation. This can lead to a 
high-impulse impedance for the static-wire system, 
thereby reducing its effectiveness. An alternative method 
could be the use of protector tubes or surge arresters on the 
power conductors in conjunction with pole grounds (37). 
Additional points about static-wire applications in mining 
are presented in chapter 13. 



300 



Impulse Performance of Ground Beds 

The impedance of a ground bed is directly connected 
with its ability to help dissipate a transient. The general 
consensus about the performance of ground beds when 
subjected to impulse currents is that the surge resistance 
of an electrode (R s ) is almost invariably less than the 
normal resistance value (R n ) as measured by an earth 
tester (9, 14, 21). Examination of oscillograph records has 
shown that under surge conditions, the ground system 
exhibits a resistance equal to or slightly higher than its 
normal value for the first 1 or 2 us. The surge resistance 
then quickly decreases to a level usually from 20% to 80% 
of the normal value (36). Figure 11.36 illustrates that as 
the value of peak surge current increases, the ratio of 
impulse to 60-Hz resistance (Rg/R n ) decreases (7). Further- 
more, ground beds that exhibit a rather high normal 
resistance show a proportionally larger resistance de- 
crease when subjected to an impulse (36). Thus, two 
ground beds with dissimilar values of R n may exhibit R s 
values that are very close to each other (8). 

Experiments have shown that in many respects soil 
behaves as a dielectric material (17). When a certain 
critical potential gradient is reached, the soil breaks down 
or arcs internally. Thus, the ground-bed resistance de- 
crease under surge conditions is due to electric discharges 
within the soil, which spark across areas of high electric- 
field intensity. As shown in figure 11.37, the presence of 
moisture in the soil acts to increase its dielectric strength 
(17). Accordingly, very dry soil breaks down much more 
readily than soil containing a minor amount of water. 
Grounding systems can be designed in such a way that 
their performance under surge conditions is optimized. 
Figure 11.38 shows how increasing the pointedness of 
electrodes results in lower values of surge resistance (21). 
As a result, driven ground rods may be superior to meshes 
in lightning prone areas since the pointed electrodes are 
conducive to high stress concentration buildup and will 
therefore cause the soil to break down more readily under 
surge conditions. 

This chapter has described the electrical transient 
and its ramifications on mine power-system safety and 
integrity. An electrical transient has been defined as the 
outward manifestation of a sudden change in circuit 
conditions. The magnitude and fast rise time of the 



1.0 

uj y 

^°2 .8 

0_UJ J— _ 

5U^ -6 



— C/l 

U_< UJ „ 

oho; .4 
Qui? 

u - IS) 





CREST IMPULSE CURRENT, kA 



overvoltage can cause damage to electrical components, 
particularly insulating materials. In addition to sound 
design practices, techniques and devices are available to 
reduce the damage done by surge voltages. The applica- 
tion of these in the design of mine power equipment is 
covered in the next two chapters. 



UJ 

< 



o 

> 



UJ 

CO 



CL 



< 

UJ 

a. 



120 


\ 4-in gap 


90 


- V ^^ 




" c ~ 1 ^^^ 


fin 


_ ?-in qnp **•*" — . 


30 


— - - 11% moisture 
i=i 0.38 % moisture 




i i i i 1 i i i i 1 i i i i 1 



5 10 

TIME TO BREAKDOWN, ^s 



15 



Figure 11.37.— Impulse breakdown of sand for two moisture 
conditions using spherical electrodes. 





2,500 












r 


A 


KEY 
No points 




2,000 




B 
C 
D 
E 


2 -cm points 
4-cm points 
6-cm points 
8-cm points 


3 

UJ 

u 

< 
H 
CO 

co 

UJ 


1,500 
1,000 

500 


\b \ 


^ 


i i 



Figure 11.36.— Ratio of impulse to 60-Hz resistance as a 
function of peak impulse current, for driven rods. 



20 40 60 

PEAK IMPULSE VOLTAGE, kV 

Figure 11.38.— Impulse characteristics of spherical elec- 
trode, with seven attached pointed protrusions of various 
lengths. 



301 



REFERENCES 

1. American National Standards Institute (New York). Stan- 
dard General Requirements for Distribution, Power, and 
Regulating Transformers. Stand. C57. 12.00-1973 (IEEE Stand. 
462-1973). 

2. . Standard for Surge Arresters for Alternating Cur- 
rent Power Circuits. Stand. C62. 1-1981 (IEEE Stand. 28-1981). 

3. . USA Standard Guide for Application of Valve-Type 

Lightning Arresters for Alternating-Current Systems. Stand. 
C62.2-1981. 

4. Stand. C84.1-1970. 

5. Armstrong, H. R. Grounding Electrode Characteristics From 
Model Tests. Trans. Am. Inst. Electr. Eng., Part 3, v. 72, Dec. 
1953. 

6. Beehler, J. E. Capacitance Switching With Power Circuit 
Breakers. Pres. at IEEE Symp. on Capacitance Switching 
Capability, New York, Jan. 31, 1968; available from IEEE, New 
York. 

7. Bellaschi, P. L. Impulse and 60 Cycle Characteristics of 
Driven Grounds. Trans. Am. Inst. Electr. Eng., v. 60, Mar. 1941. 

8. Bellaschi, P. L., R. E. Armstrong, and A. E. Snowden. Im- 
pulse and 60 Cycle Characteristics of Driven Grounds-II. Trans. 
Am. Inst. Electr. Eng., v. 61, Dec. 1942. 

9. Bewley, L. V. Theory and Tests of the Counterpoise. Trans. 
Am. Inst. Electr. Eng., v. 53, Aug. 1934. 

10. Boehne, E. W. Energization Surges of Capacitive Circuits. 
IEEE Winter Power Meet., Conf. Paper 70CP235-PWR, 1970. 

11. Buss, E. W., R. C. Dugan, and P. C. Lyons. Vacuum Circuit 
Breakers and Dry-Type Transformers Special Considerations for 
Mining Operations. Paper in Conference Record- IAS 12th Annual 
Meeting (Los Angeles, CA, Oct. 1977). IEEE, 1977. 

12. Card, R. H. Earth Resistivity and Geological Structure. 
Trans. Am. Inst. Electr. Eng., v. 54, Nov. 1935. 

13. Coyle, M. Effective Protection Schemes Smooth Transient 
Pains. Electron. Prod. Mag., v. 18, May 1976. 

14. Davis, R., and J. E. M. Johnston. The Surge Characteristics 
of Tower and Tower-Footing Impedances. J. Inst. Electr. Eng. 
(London), v. 88, Oct. 1941. 

15. Dornetto, L. D. The Importance of Grounding Systems in 
the Protection of Personnel and Equipment. Paper in Mine Power 
Distribution. Proceedings: Bureau of Mines Technology Transfer 
Seminar, Pittsburgh, Pa., March 19, 1975. BuMinesIC 8694, 1975. 

16. Eaton, J. R. Electrical Power Transmission Systems. 
Prentice-Hall, 1972. 

17. Impulse Characteristics of Electrical Connections to 

the Earth. Gen. Electr. Rev., v. 47, Oct. 1944. 

18. General Electric Co. (Syracuse, NY). Transient Voltage 
Suppression Manual. 1976. 

19. Greenwood, A. N. Electrical Transients in Power Systems. 
Wiley, 1971. 

20. Greenwood, A. N., D. R. Kurtz, and J. C. Sofianeko. A Guide 
to the Application of Vacuum Circuit Breakers. IEEE Trans. 
Power Appar. and Syst, v. 90, July/ Aug. 1971. 



21. Hemstreet, J. G., W. W. Lewis, and C. M. Foust. Study of 
Driven Rods and Counterpoise Wires in High-Resistance Soil on 
Consumers Power Company 140-kV System. Trans. Am. Inst. 
Electr. Eng., v. 61, Sept. 1942. 

22. Institute of Electrical and Electronics Engineers (New 
York). Recommended Practice for Electric Power Distribution for 
Industrial Plants. Stand. 141-1986. 

23. Recommended Practice for Grounding of Industrial 

and Commercial Power Systems. Stand. 142-1972. 

24. Kunjara, N. A., G. P. Russell, and E. K. Stanek. Prediction 
and Suppression of Electrical Transients in Mine Electrical 
Systems. Paper in Conference Record -IAS 12th Annual Meeting 
(Los Angeles, CA, Oct. 1977). IEEE, 1977. 

25. Lewis, W. W. The Protection of Transmission Systems 
Against Lightning. Wiley, 1950. 

26. Long, R. J. The Selection and Application of Vacuum Circuit 
Breakers for Open Pit Mining and Excavation. Paper in Con- 
ference Record- IAS 11th Annual Meeting (Chicago, IL, Oct. 
1976). IEEE, 1976. 

27. Marshall, J. L. Lightning Protection. Wiley, 1973. 

28. McCann, G. D., and D. E. Morgan. Field Disturbances Pro- 
duced by Lightning. Trans. Am. Inst. Electr. Eng., v. 62, 1943. 

29. McEachron, K. B. Multiple Lightning Strokes. Trans. Am. 
Inst. Electr. Eng., v. 53, 1934. 

30. Morley, L. A., F. C. Trutt, and J. L. Kohler. Final 
Report -Evaluation of Coal-Mine Electrical-System Safety (grant 
0155003, PA State Univ.). BuMines OFR 160-81, 1981; NTIS PB 
82-139338. 

31. Ohio Brass Co. (Mansfield, OH). How Does a Distribution 
Class Surge Arrester Work? Tech. Book. 2470-H, undated. 

32. Pflantz, H. M. Analysis of Multiple Prestrike and Interrup- 
tion Phenomena in Capacitive Circuits. Pres. at IEEE Summer 
Meet, of Power Eng. Soc, San Francisco, CA, July 1972; available 
from IEEE, New York. 

33. Rudenberg, R. Transient Performance of Electric Power 
Systems. MIT Press, 1967. 

34. Smith, E. P. Lightning Arrester Applications for Mine 
Power Systems. Paper in Conference Record-IAS 11th Annual 
Meeting (Chicago, IL, Oct. 1976). IEEE, 1976. 

35. Titus, C. H. Evaluation and Feasibility Study of Isolated 
Electrical Distribution Systems in Underground Coal Mines. Apr. 
1972; NTIS PB 213 741. 

36. Towne, H. M. Impulse Characteristics of Driven Ground. 
Gen. Electr. Rev., v. 31, Nov. 1928. 

37. Westinghouse Electric Corp. (East Pittsburgh, PA). Elec- 
trical Transmission and Distribution Reference Book. 4th ed., 1964. 

38. Westinghouse Electric Corp., Relay -Instrument Div. 
(Newark, NJ). System Neutral Grounding and Ground Fault Pro- 
tection. PRSC-4E, Ind. and Commer. Power Syst. Appl. Ser., Feb. 
1986. 



302 



CHAPTER 12.— MINE POWER CENTERS 



The major power equipment in mining power centers, 
switchhouses and substations, was introduced in chapter 
1, and an elementary overview of the protective circuitry 
involved was given in chapter 9. Other chapters have 
added numerous basic concepts and techniques whose 
thrust has actually provided the background for this 
chapter and the next, where the major power equipment 
used in mines is discussed in detail. In this chapter, the 
internal components and construction of typical mine 
power centers are covered. It should be noted that no 
formal standards have yet been developed for this equip- 
ment; hence, extensive use is made of information pro- 
vided by major manufacturers serving the industry and by 
several mine operators. 

Many circuit diagrams in this chapter apply directly 
to underground coal mining. With a few notable excep- 
tions, underground power equipment is the most complex 
electrical equipment found in mining. An understanding 
of this equipment should, therefore, lead to comprehension 
of that used in any other mining system. In the para- 
graphs that follow, some material contained in preceding 
chapters is reviewed, but usually the content is either 
changed or expanded. The main reason is to enhance 
many section presentations without requiring reference to 
other information. 



EQUIPMENT SPECIFICATIONS 

An important objective of this chapter and the next is 
to provide sufficient information for an individual to draw 
up the detailed specifications required when procuring a 
piece of power equipment to fit a mining need. It is often 
stated that a manufacturer supplies what the customer 
requests. Since there are no official standards nor recom- 
mended practices at present, and given the variability of 
most equipment, every piece of mine power apparatus 
must be specified individually. If these specifications are 
not drawn up in complete detail, down to each nut and 
bolt, the purchaser might not receive what he or she 
intended. 

Making detailed equipment specifications does not 
imply criticism of equipment manufacturers. It is a fact 
that many companies employ very capable applications- 
oriented engineers who have the total respect of the 
mining industry. Some can deliver precisely matched 
power systems just by knowing the mining equipment 
being used and the seam in which it is operating. Even in 
these cases, however, a complete specification from the 
buyer is still wise in terms of the protection it affords both 
the manufacturer and the customer. In situations where 
manufacturers receive specifications they believe are 
faulty or feel money can be saved by another approach, 
most manufacturers will contact the buyer for clarification 



rather than following the specifications rigidly and with- 
out question. 

A listing of the minimum number of items recom- 
mended for a specification is provided below. 



1.0. 
2.0. 
3.0. 



4.0. 



5.0. 



6.0. 



7.0. 



8.0. 



Time Schedule 
Work by Purchaser 
General 

3.1. Intended use of equipment 

3.2. Enclosure specifications 

3.3. Special needs 
Internal Components 

4.1. One-line diagram 

4.2. Description of each component 
Drawing and Manual Requirements 

5.1. With manufacturer's bid 

5.2. Before construction starts 

5.3. With equipment delivery 
Inspections 

6.1. During construction and prior to delivery, 

by purchaser 

6.2. Manufacturer compliance with local, 

State, and Federal regulations applicable 
to company 

6.3. Equipment (being built) compliance with 

applicable local, State, and Federal regu- 
lations 
Guarantee 

7.1. Minimum guarantee or warrantees de- 

manded by buyer 

7.2. Request for proper storage requirements 

before service 
Delivery Dates 






1 The author wishes to thank Thomas Novak, associate professor of 
mineral engineering, the University of Alabama, who prepared the origi- 
nal manuscript for this chapter while he was an instructor in mining 
engineering at The Pennsylvania State University. 



A document prepared to this format could be used as part 
of a purchase order directly to one manufacturer or as a 
solicitation for bids from several companies. Usually, such 
information is preceded by a general title, a phrase ex- 
pressing what the specification covers, and the number of 
units desired. The next paragraphs will present sugges- 
tions for the content within each item of the document. 
The individual making the specification is referred to as 
the buyer or purchaser. 

The time schedule alerts the manufacturer to when 
the buyer expects work on the equipment to begin. Fur- 
thermore, it serves notice that the manufacturer must 
have enough personnel, equipment, and material to com- 
plete work within the delivery time stated. (Obviously, if 
the number of sources for the equipment is restricted— and 
sometimes there could be only one, the buyer must accept 
the seller's delivery dates.) Item 2.0 should contain a 
statement that the purchaser will accept only equipment 
that meets the specifications. In other words, the units 
may be rejected if they do not comply, and the purchaser 
has no obligation to pay. 

A detailed account of the intended use for the equip- 
ment should be given under item 3.1 so the manufacturer 
knows exactly what the specification covers and what the 
buyer expects to do with the unit after receiving it. Item 



303 



3.2 primarily applies (but is not restricted) to unit- 
designed equipment. Content examples include: 

• Minimum material requirements. 

• Physical protection of internal parts. 

• Desired mounting (tire, skid, or rail). 

• Measures to prevent water from entering enclosure. 

• Maximum length, width, and height. 

Minimum materials should always be stated when the 
specification is going out for bid and when ruggedness for 
the mining environment is imperative. To prevent any 
inadequate bids, it is wise to provide minimum acceptable 
steel thicknesses for the base plate, frame, end plates, and 
covers. Specification of maximum dimensions is essential 
for underground operations. It may seem absurd, but there 
have been many instances where equipment has been 
purchased that is too large to fit down a shaft or slope. 
Special needs, item 3.3., refers to such things as engraved 
nameplates for all major components and labeling of 
internal wiring and terminal blocks. 

Item 4.0, internal components, is commonly the larg- 
est part of a specification. It should provide a complete 
listing of all components with one-line diagrams to show 
how they are to be connected. It is also helpful to state how 
each component is to be used and why each is included. 
Special needs such as transformer ratios, trip settings, 
and minimum insulation levels should be given where 
applicable. 

Item 5.0 is concerned with requirements for drawings 
and manuals. If the specification is being sent out for bids, 
it is good to request both one-line diagrams and schemat- 
ics of physical layout from prospective manufacturers. 
With unit-designed equipment, details for physical layout 
would include the basic frame, the mounting arrange- 
ment, and major component layouts. The manufacturer 
should be asked to submit all applicable outlines, arrange- 
ments, and schematics for approval prior to starting con- 
struction. The specification should also state the number 
of parts manuals and instruction manuals to be supplied 
on delivery of the equipment. Drawings, including some of 
reproducible quality, should also be requested where 
applicable. 

Item 6.0 should be included to protect the customer. 
The first entry must relate that the buyer has the right to 
inspect the equipment during manufacture and prior to 
delivery and that any discrepancies from the specification 
must be corrected. At times, orders can be construed as 
contracts; therefore, the buyer might be held (in part) 
legally responsible if the manufacturer violates any local, 
State or Federal ordinances, codes, acts, regulations, or 
laws. Thus, the specification must demand compliance 
with all applicable statutes; this should be extended to any 
subcontractors or suppliers that the manufacturer might 
use. The last entry for item 6.0 is a statement that the 
specified equipment shall comply with all Federal and 
State regulations and requirements that apply to the 
intended use in the State in which it will operate. 

Guarantees or warranties (item 7.0) cannot be ob- 
tained on some mine power equipment. Regardless, a 
request for a minimum guarantee is wise, for example, a 
1-yr guarantee that begins after the equipment is placed 



in operation. The manufacturer should also be asked to 
supply any special instructions or precautions necessary 
for proper equipment storage prior to its being placed in 
service. 

Finally, the specification must relate the dates that 
the equipment is to be delivered to the mining operation. 

Establishing an adequate specification is not always 
an easy process. As will be seen, certain types of mine 
power equipment could require a document of numerous 
pages. However, a good specification might be the differ- 
ence between having a unit that performs superbly 
throughout the life of the mine or one that is a headache 
from the day it is placed in service. 



MINE POWER CENTERS 

The power or load center is one of the most essential 
power-system units for underground mines and, in a 
simpler form, for surface mines. Its primary function is to 
convert the distribution voltage to utilization voltage for 
operating equipment throughout the mine. It must incor- 
porate protective circuitry to ensure safe, efficient, and 
reliable operation. In effect, the power center could be 
considered a portable substation, but because of the ways 
in which it is used, its main component (and perhaps the 
enitre unit) might also be classified as a distribution 
transformer. 

The electrical components of the power center are 
usually metal clad, that is, housed in a heavy-duty steel 
enclosure that may be tire mounted, skid mounted, or 
track mounted. Illustrations of typical underground coal 
mining units are given in figure 12.1 (9, 12). 2 Towing lugs 
or pin-and-link couplers are commonly provided on each 
end of the enclosures to permit towing as mining advances 
or retreats. Bumpers or check plates are often installed to 
protect externally mounted components, such as couplers, 
from damage by mobile equipment. Similar enclosures can 
be used in surface mines, but here the simplest power 
center consists of outdoor components assembled on a 
flatbed trailer with a fence or gates to discourage unau- 
thorized entry. 

There is no such thing as a "standard" power center 
because of the variety of mining practices and the numer- 
ous types of mining equipment used. The power center 
may supply only 1 motor or as many as 20 pieces of 
machinery; it may be totally ac, dc, or a combination of ac 
and dc. The distribution voltage received by the power 
center can be 4.16, 7.2, 12.47, 13.2, or 13.8 kV. The 
outgoing ac utilization voltage may be 480, 600, 995, 1,040 
V, or a combination of 995 or 1,040 V with one of the lower 
voltages. Dc can be at 300 or 600 V, but is almost always 
300 V for face applications. As a result of this variety, 
manufacturers custom-build the units to meet the individ- 
ual needs and specifications of the customer. However, 
there is a general design philosophy central to all power- 
center types and this forms the basis of the following 
discussion of a typical ac power center and a combination 
ac-dc power center. 



2 Italicized numbers in parentheses refer to items in the list of references 
at the end of this chapter. 



304 




Tire mounted 



The major electrical components of a typical mine 
power center are shown in the schematic of figure 12.2, 
and a possible component placement is provided in figure 
12.3. The circled numbers in figure 12.2 will be used to 
indicate component location with respect to the overall 
system throughout the following descriptions of individual 
components. 




Skid mounted 




Track mounted 

Figure 12.1.— Typical power centers used in underground 
coal mines. 



HIGH-VOLTAGE CABLE COUPLER 

Incoming power is usually supplied to the center from 
the distribution cable through a high-voltage cable cou- 
pler (fig. 12.2, No. 1). The receptacle (typically with female 
contacts) is cable mounted, while the plug (male contacts) 
is gear mounted. Although the conductor pins are recessed 
in the coupler housing, dust caps should be provided for 
each side of the coupler pairs to protect the contacts when 
disengaged. 

When being disconnected, the coupler is designed so 
that the pilot contact is broken first, the line conductors 
second, and the grounding conductor last. As was dis- 
cussed in chapter 9, the pilot contact is interlocked with 
the upstream high-voltage circuit breaker, which protects 
the incoming line. If the incoming power is energized, the 
power will be tripped by the associated ground-check 
monitor prior to breaking the power contacts of the cou- 
pler. Having the grounding contact break last ensures that 
the frame is tied to earth ground whenever the power 
center is energized. 

A feedthrough receptacle may be provided as shown in 
figure 12.4; the contacts are typically female. This permits 




P 



o->o — n — n- 



f® 







HHHHHHHH h— j_ 





ff^T 



5 




KEY 

1 High -voltage coupler 

2 Interlock switches 

3 Emergency-stop switch 

4 Disconnect switch 

5 Pilot-break monitor 

6 High-voltage fuses 

7 Surge arresters 

8 Surge capacitors 

9 Power transformer 

10 Temperature device 

1 1 Grounding resistor 

12 Busway 

13 Outgoing circuit breaker 

14 Main circuit breaker 

15 Voltage metering 

16 Current metering 

17 Outgoing cable coupler 



Figure 12.2.— Schematic illustrating major components in power center. 



305 



Outgoing 

cdbl 



LOW-VOLTAGE TRANSFORMER 
COMPARTMENT COMPARTMENT 



HIGH-VOLTAGE 
COMPARTMENT 



High-voltage 
coupler 




if 1 Feedthrough 
1 J receptacle 



Figure 12.3.— Top view of mine power center showing place- 
ment of many internal components. 




To 
id-bn 
switch 



\ lo 

"X> load-break 

■^ Qwil 



Figure 12.4.— Interconnections between input and feed- 
through receptacles. 



distribution power to be supplied (or continued) through 
the power center at other higb-voltage loads. A dust cap is 
again provided for use when the feedthrough receptacle is 
not in service. This cap also shorts the pilot contact to the 
grounding contact to provide a closed path for the ground- 
check monitor interlocking with the upstream circuit 
breaker. The conductors between the input and feed- 
through receptacles should be sized to the maximum 
capacity of the distribution system. 



For added safety, some manufacturers sectionalize 
their power centers into three separate compartments: 
high voltage, transformer, and low or medium voltage (fig. 
12.3). If the equipment has no barriers, interlock switches 
should be placed on the low-voltage or medium-voltage 
exterior covers, with their contacts in series with the 
incoming pilot circuit. These additional switches are rec- 
ommended even with the compartment segregation, but 
then the interlocks are connected to trip circuit breakers 
associated with the transformer secondary, thus avoiding 
tripping the upstream switchhouse. 

An emergency-stop button (fig. 12.2, No. 3) should also 
be provided, and its function is similar to the interlock. It 
consists of a normally closed {NO set of contacts in series 
with the interlock switches. If the stop button is de- 
pressed, its contacts are opened, opening the incoming 
pilot circuit, and thus tripping the upstream circuit 
breaker. The switch should not automatically reset to the 
normally closed position after being depressed; manual 
resetting of the switch should be required. 



DISCONNECT SWITCH 

The disconnect or load-break switch (fig. 12.2, No. 4) is 
a mechanically operated air-type switch whose primary 
function is to allow a quick means of disconnecting the 
primary of the power-center transformer. A spring-loaded 
or torsion-bar mechanism provides the quick-make and 
quick-break operation, which is independent of the speed 
of the manually activated handle. Observation windows 
are provided in the power-center enclosure, and the switch 
can serve as a visual disconnect. 

Disconnect switches have no interrupting capability, 
but load-break switches do. Load-break switches are able 
to interrupt currents that are not in excess of their 
continuous-current rating. They are also rated for a max- 
imum interrupting capacity, but at this level they are 
designed for one-time interruption only. Load-break 
switches are more expensive than disconnects, yet many 
mine power engineers prefer them because of their extra 
ruggedness. Some States require their use. Continuous- 
current ratings of 400 and 600 A and a 15-kV voltage 
rating are common for mining applications. Table 12.1 
illustrates typical ratings of a 400-A load-break switch. 



Table 12.1.— Typical current ratings of 400-A load-break 
switch 



INTERLOCK SWITCHES 

Electrical interlock switches (fig. 12.2, No. 2) are 
presently used in the high-voltage and transformer com- 
partments of mine power equipment. Strategically located 
around exterior top and side protective covers, they act to 
deenergize the internal power circuitry if the panels are 
removed, and thus help to prevent accidents caused by a 
worker's contacting energized components. The normally 
open (NO) contacts of the switches are connected in the 
pilot circuit of the incoming distribution cable as shown in 
figure 12.2. With covers in place, the interlock switches 
are depressed and their associated contacts are closed. 
This provides a closed path for the pilot circuit. When a 
cover is removed, the switch contacts are opened, thus 
causing the upstream circuit breaker to trip because of the 
loss of continuity in the pilot circuit. 



Full-load current 

Rated interrupting capacity 

Maximum interrupting capacity. 

Short-time current (1 s) 

Impulse current 

Close-and-latch current 



Rating, A 

400 

400 

1,200 

20,000 

45,000 

37,000 



The only means of activating a load-break switch 
should be by the manually operated handle accessible 
from outside the load-center enclosure. Its operating mech- 
anism should not be tied into the power-center protective 
circuitry. More will be said about this recommendation 
soon. 

If a disconnect switch is allowed and employed, a 
pilot-break monitor (fig. 12.2, No. 5) is required to inter- 
lock the switch handle with the pilot circuit of the incom- 
ing distribution system. This allows the upstream circuit 



306 



breaker to interrupt the circuit prior to opening the switch 
contacts. It is also desirable to use pilot interlocking on 
load-break switches to extend the life of the contacts. 
Again, some States require both load-break switches and 
pilot -break monitoring in order to minimize any danger of 
operating the switch under load. The balance of this 
chapter will use or imply the term load-break for all 
disconnect-switch applications. 



HIGH-VOLTAGE FUSES 

A fuse has been defined as an overcurrent protective 
device with a circuit-opening fusible element that is 
heated and severed by the passage of current through it. 
Current-limiting fuses (fig. 12.2, No. 6) are the type used 
in mine power centers. Their main purpose is to protect 
the high-voltage section, particularly the transformer, 
during short circuits. High-voltage circuit breakers are not 
recommended here, because of their potential for creating 
transients. 

For resistance-grounded mine power systems, the fuse 
voltage rating should be based on the line-to-line voltage. 
The following two rules-of-thumb are recommended for 
determining the fuse current rating that will ensure that 
fuse action will not be triggered by the transformer inrush 
current. 

1. The fuse should be able to withstand 12 times the 
rated current of the transformer for 0.1 s without element 
damage. 

2. The element should be able to withstand 25 times 
the rated current of the transformer for 1/2 cycle. 

Based upon the above criteria, the continuous-current 
rating of the current-limiting fuses usually falls in the 
range of 1.5 to 2.5 times the rated transformer current. 

Transient overvoltages can be generated during the 
operation of the fuses. The magnitude of the overvoltage 
depends upon the point of the waveform at which initia- 
tion of the fault occurs and the size and design of the fuse 
element. For ribbon-element fuses, the generated peak arc 
voltage is primarily a function of the system voltage, as 
shown in figure 12.5 (10). The peak arc voltage should be 
compared with the minimum 60-Hz sparkover level of the 
surge arresters. The sparkover level of the arresters^ 
discussed in the next section, should be multiplied by V2 
to obtain a peak voltage. If this value is greater than the 
peak arc voltage, arrester sparkover will not result from 
fuse interruption. 

The fuses are sometimes used in conjunction with an 
automatic (fused) load-break switch, which has the objec- 
tive of preventing the power center from operating under a 
single-phase condition if only one fuse blows. The fuses 
normally have actuators located at one end of their cas- 
ings. In the event that a fuse element is blown, the 
actuator extends outward from the fuse end and activates 
the trip mechanism of the switch. This causes the switch 
to open and eliminates the single-phase condition. How- 
ever, for a three-phase fault, the possibility exists that one 
fuse may blow before the others, which could result in the 
load-break switch attempting to clear a fault current 
higher than its designated rating. As a result, this type of 
automatic load-break switch is not recommended in power- 
center applications. 

Single-phasing protection is better accomplished at 
the secondary bus, using relays to monitor for a single- 



< 

b 

o 
> 

o 

ir 
< 

< 

UJ 
Q. 



X 

< 



uu 
90 

80 
70 














/ 














/ 
















60 
50 
40 












































30 
















?0 
















10 































5 10 15 20 25 30 
CIRCUIT VOLTAGE, kV 



35 



Figure 12.5.— Graph illustrating transient crest voltage 
caused by ribbon-element current-limiting fuse operation. 



phase or phase-reversal condition. The relay contacts can 
be employed to activate the undervoltage release or shunt 
trip of the main breaker on the secondary. If a main 
breaker is not used, the contacts can be connected directly 
at the output of the control transformer secondary wind- 
ing, so all outgoing circuit breakers will trip through their 
undervoltage releases if the relay is actuated. 



SURGE ARRESTERS 

The surge arrester (fig. 12.2, No. 7) is a device de- 
signed to limit dangerous transient overvoltages to safe 
levels. Lightning, switching surges, and some faults result 
in transient overvoltages that can exceed the insulation 
levels of power-system equipment. Since it is not always 
economically possible to rate the equipment insulation 
above the surges, overvoltages must be clamped or sup- 
pressed to tolerable levels. The function of an arrester is 

• To discharge the energy associated with a transient 
overvoltage; 

• To limit and interrupt the 60-Hz current that fol- 
lows the transient current through the arrester; and 

• To return to an insulating state without interrupt- 
ing the supply of power to the load. 

Valve-surge arresters of the low-sparkover distribution 
class are used almost exclusively in mine power centers. A 
complete discussion of these devices as well as application 
tables was provided in chapter 11. 

The voltage rating of a surge arrester is the highest 
power-frequency voltage at which the arrester is designed 
to operate. As mine power systems utilize resistance 
grounding and fall into the category of noneffectively 
grounded systems, the surge-arrester voltage rating must 
be selected on a line-to-line basis. However, the rating 
should be 5% to 10% above the nominal distribution 
voltage to allow for voltage-regulation compensation. 



307 



The component protected by the surge arrester is the 
main power transformer. The key to protection is coordina- 
tion between the transformer insulation withstand, or BEL 
(basic impulse insulation level), and the characteristics of the 
transient voltage the arrester lets through. Dry transformers 
are used almost exclusively in present-day power centers, 
and their insulation strength does not increase significantly 
above their BIL as the duration of the applied pulse de- 
creases. Therefore, the voltage-time withstand characteristic 
can be plotted as a flat line with a value equal to the BEL, as 
shown in figure 12.6 (11). The arrester characteristic, includ- 
ing impulse sparkover and discharge voltage, is also illus- 
trated. The margin of protection is the difference between 
the transformer BEL and the arrester discharge characteris- 
tic at any given instant of time. This value should not be less 
than 20% if a 5,000-A surge discharge current is assumed. 
Note that the worst case discharge of 20,000 A need not be 
used in typical power-center locations. 

The physical location of surge arresters within the 
power center is important for effective operation. The two 
critical distances are the conductor lengths to the line 
conductor and to the power-center frame ground, and the 
distance between the arrester and the transformer. To 
maximize arrester performance, the arrester leads must 
be as short and straight as possible and made of No. 6 
AWG solid copper or larger. Also, the surge arrester and 
arrester connections to the line conductors should be as 
close to the transformer primary terminals as possible. 
Preferably, the surge arrester leads should be connected 
right at these terminals. 

Surge capacitors (fig. 12.2, No. 8) can be found in 
numerous mine power centers. Their intended purpose is 
to lessen the severity of transient overvoltages caused by 
current chopping in vacuum circuit breakers. Their effec- 
tiveness is based upon the ability to limit the rate-of-rise of 
transient overvoltages and to reduce the system character- 
istic impedance. However, as was discussed in chapter 11, 
surge capacitor use in load centers is generally not re- 
quired, provided that the surge arresters are correctly 
applied, and a minimum practical cable length (100 ft) 
exists between the power center and the upstream circuit 
breaker. 



BIL (full-wave withstand) 



< 



o 
> 



Minimum margin 
of protection 



. Front-of-wave 
[ sparkover 




■*- 



Discharge-voltage 

maximum characteristics 

of arrester 



Impulse voltage rising 

100 kV/^s for each 

12 kV of arrester rating 



1 2 3 4 5 6 

TIMERS 

Figure 12.6.— Comparison of transformer withstand 
characteristic and surge arrester withstand characteristic. 



TRANSFORMERS 

The main transformer (fig. 12.2, No. 9) can be consid- 
ered the heart of the power center, since its primary 
function is to convert the distribution voltage to utiliza- 
tion. Proper selection is therefore imperative from the 
standpoint of safety, efficiency, and reliability. Fortu- 
nately, the IEEE (2, 4) has established classifications and 
specifications for determining transformer characteristics 
as an aid in design and application, and these regulations 
are in general use in the mining industry. 

The IEEE uses three items for the general classifica- 
tion of transformers: distribution or power, insulation, and 
substation or unit substation. The capacity or kilovoltam- 
pere rating determines whether the transformer is classed 
as a distribution or power unit. Distribution transformers 
fall in the range from 3 to 500 kVA, with power transform- 
ers having capacities greater than 500 kVA. Power-center 
transformers can be of either classification, since they may 
range from 150 kVA for a rectifier or belt drive, up to 2,250 
kVA for a longwall section. Capacities are rarely greater 
than 1,250 kVA. 

A transformer may be further classified by the insu- 
lation system as liquid or dry. Liquid insulation includes 
mineral oil or synthetic fluids, and dry transformers are 
ventilated or sealed gas-filled types. Ventilated dry units 
are used almost exclusively in mine power centers. With 
convection cooling and air insulation, dry transformers 
have the following advantages: 

• Toxic gases cannot be released. 

• The transformer cannot explode or catch fire. 

• There is no oil or other liquid to spill, leak, or 
dispose of (a critical problem with PCB, poly chlorinated 
biphenyls). 

• The transformer is virtually maintenance-free be- 
cause there are no valves, pumps, or gauges. 

The substation-transformer classification refers to direct 
or overhead-line termination facilities, while a unit- 
substation transformer has an integral connection to pri- 
mary or secondary switchgear. Primary refers to a voltage 
of 1,000 V or higher, while secondary refers to a rating less 
than 1,000 V. For practical purposes, 1,040 V is also 
considered 1,000 V. Therefore, a load-center transformer 
can be considered a unit-substation class. 

Specifications 

The IEEE transformer specifications (4) that apply to 
mine power centers include 

• Capacity or kilovoltampere rating, 

• Phases, 

• Frequency, 

• Voltage rating, 

• Voltage taps, 

• Winding connections, 

• Impedance, 

• BIL, 

• Temperature rise and insulation, 

• Cooling. 

Each will be discussed in general terms in the next para- 
graphs. There are a few exceptions to the IEEE standard. 



308 



There is no set formula for determining the kilovolt- 
ampere rating for a power-center transformer. For con- 
stant loads, it is a common rule-of-thumb to allow 1 kVA 
for each horsepower of connected load. However, the min- 
ing process does not produce a constant load (that is, all 
connected motors are not operating at the same time on a 
continuous basis); hence, using the rule-of-thumb will 
normally result in oversizing the transformer. Past expe- 
rience and demand factors established by manufacturers 
and operators, along with the horsepower of the connected 
load, are essential for determining the transformer capac- 
ity. For typical underground mining sections, the kilovolt- 
ampere rating may lie within the range of 50% to 80% of 
the connected horsepower. (Chapters 4 and 8 contain 
additional information on demand factors.) 

As an example of sizing, consider the following 
continuous-mining section: 

1. 370-hp continuous miner (gathering head, pump, 
tram, and cutter motors), 

2. Two 125-hp shuttle cars (traction, conveyors, and 
pump motors), 

3. 80-hp roof bolter (traction and pump motors), and 

4. 50 hp of auxiliary equipment (section fan, sump 
pump, hand tools, etc.). 

The total connected horsepower sums to 750 hp. If the 
demand factor has been determined to be 60%, the effec- 
tive load would be 450 kVA, and a 500-kVA transformer 
would be selected. However, flexibility must always be 
considered in mining applications, and as a result, it may 
be necessary to select the next higher kilovoltampere 
rating to accommodate anticipated additional loads. 

Over the years, power-center manufacturers have ar- 
rived at certain standard or typical transformer capacities, 
created mainly by repeat demand of the industry. These 
are- 
Ac output only (section may have outboard rectifier for 
dc load): 150, 225, 300, 500, 600, 750, 1,000, and 1,250 
kVA. 

Dc only (often termed rectifiers; rating of unit com- 
monly in kilowatts): 100, 150, 200, 300, 500, and 750 kVA. 

Combination ac-dc (typical dc capacity; on larger 
units, 200 and 300 kW for dc also available): 300/150, 
500/150, 600/150, 750/150, 1,000/150, and 1,250/150 kVA. 
Many manufacturers will supply capacities at any incre- 
ment of 50 kVA up to 2,250 kVA, the maximum allowed by 
Federal regulations for face applications in underground 
coal mines. 

Three-winding transformers are necessary in many 
power centers, for example, when ac face equipment has 
two rated voltages such as a 950-V continuous miner 
working with 550-V machines, or in the common applica- 
tion of mixed ac and dc machinery. In these cases, the 
capacity of each transformer winding (primary, secondary, 
and tertiary) must be individually rated. 

All mine power-center transformers are three phase, 
being either three single-phase units where each trans- 
former is rated at one-third of the total required capacity, 
or integral three-phase types with construction allowing 
field replacement of failed windings. The integral three- 
phase transformers are preferred because they have equal 
reliability, lower cost, and higher efficiency, take less 
space, and have fewer exposed interconnections than three 
single-phase units. 

Transformers should be rated at the standard com- 
mercial frequency in the United States, which is 60 Hz. If 



a transformer rated at 60 Hz is applied to a 50-Hz system, 
the voltage and the kilovoltampere rating must be reduced 
by 20%. Transformers rated at 50 Hz can be operated at 60 
Hz, but the efficiency and regulation are reduced, since the 
reactance is directly proportional to the frequency. 

The primary voltage rating of the transformer is 
dictated by the distribution voltage (that is, 4.16, 7.2, or 
12.47 kV, and so on). The voltage rating of the secondary 
winding must match the voltage of the utilization equip- 
ment. In practice, the transformer secondary should be 
rated at a higher voltage (10% is common) than shown on 
the nameplate of the utilization equipment, to allow for a 
voltage drop in the trailing cables. The secondary winding 
of power-center transformers is most commonly rated at 
480 or 600 V, relating to 440- or 550-V equipment, respec- 
tively. Secondaries supplying 950-V machines are usually 
rated at 1,040 V, except for machines that cannot be 
remotely controlled or in States where face voltages over 
1,000 V are not allowed. In these instances, 995 V is often 
specified. The transformer voltage for dc applications is 
dictated by the rectifiers and will be presented in a 
separate section. 

Power-center transformers should be provided with 
voltage taps on the primary windings to account for 
voltage fluctuations and line losses in the distribution 
system. Five fully-rated taps at 2.5% increments are 
usually available, with the middle tap being rated at the 
nominal distribution voltage. Figure 12.7 illustrates this 
arrangement for a 7.2-kV primary on a per-phase basis. 

Delta primary and wye secondary are the preferred 
connections for standard two-winding power transformers 
and are the connections commonly used in mine power 
centers. An external busing from the neutral of the wye 
secondary provides an easy means for resistance grounding, 
as required by Federal regulations. The delta-connected 
primary provides isolation of the distribution circuit from 
the utilization circuit with respect to zero-sequence currents 
resulting from exciting currents or secondary ground faults. 
The delta-wye connection also stabilizes the secondary neu- 
tral point and minimizes the production of third-harmonic 
voltages. Additional information concerning the reasoning 
behind using delta-wye transformers can be found in chap- 
ters 4 and 9, including the reasons wye-wye transformers are 
not recommended. 

In certain situations, a delta-connected secondary 
may be specified or required, and the primary may be 
delta or wye in this case. If neutral grounding is desired or 
required, a grounding transformer is needed to derive a 
neutral. The zig-zag grounding transformer, illustrated in 
figure 12.8, is sized according to the continuous-current 




SECONDARY 



Figure 12.7.— Typical primary winding taps on power cable 
transformer. 



309 



rating of the neutral-grounding resistor. Standard single- 
phase or integral three-phase transformers, arranged in a 
wye-delta configuration, can be used as grounding trans- 
formers in lieu of the special zig-zag type, as shown in 
figure 12.9. 

The transformer impedance is sometimes referred to 
an -impedance voltage because of the common means of 
measurement. Impedance voltage is the voltage required 
to circulate rated current through one of two windings of a 
transformer when the other winding is short-circuited, 
with the windings connected as for rated voltage operation 
(4). 

An example will help to illustrate this procedure. 
Figure 12.10 shows a three-phase delta- wye distribution 
transformer with the indicated ratings. The secondary of 
the transformer is short-circuited while a three-phase 
variable voltage source is connected to the primary. The 
voltage source is wye connected so the impedance voltage, 
Vj, will be a line-to-neutral voltage. The source voltage is 
increased until the ammeter reads the rated or full-load 
current of the transformer, I FL , which can be determined 
by 



W = 



s 

V3 V 



(12.1) 



where I FL = full-load current of transformer, A, 

S = rated transformer capacity, VA, 
and V = rated voltage at transformer winding termi- 
nals, V, 

or for the primary, 




Line leads 




Neutral leads 

Windings on coil Schematic 

Figure 12.8.— Zig-zig grounding transformer. 



To line conductors 

A A 



ILUU 



UuuJ 



Luuu 



Neutral 

grounding 

resistor 



I 



Figure 12.9.— Delta-wye connection for deriving a neutral. 



J-TTT — 



500,000 
V3 (7200) 



= 40.1 A. 



Consider the case where V ; = 250 V, when the ammeter 
reads 40.1 A. The impedance voltage is normally ex- 
pressed in terms of a per-unit value; therefore, 

250 V nn „ 

V - = 4160V = 006 P u ' 



The line-to-neutral rating of the transformer (4,160 V) is 
used since the impedance is being measured as a line- 
to-neutral potential. For this example, the line-to-neutral 
impedance of the transformer, with all quantities reflected 
to the primary, would be 



LiTy 



V: 250 V 



I, 



40.1A 



= 6.23 Q. 




7, 200 /480 V 
500 kVA 

Figure 12.10.— Technique for measuring transformer im- 
pedance. 



The transformer impedance can be measured in terms of a 
per-unit quantity by using the ratings of the transformer 
as base values, as follows: 



kVA b = 500 kVA, 



kV b = 



I b = 



7,200 

~7T 



kVA 



= 4.16 kV (line to neutral), 



500 



V3kV b " V3 (7.2) " 4(U A ' 



_ V b _ 4,160 
b I b "" 40.1 



= 103.7 0. 



Thus, the per-unit value of the transformer impedance is 
Z„ 



6.23 nn „ 

^" ^ To3?Tfi = 006 pu - 



This yields the same results; therefore in terms of per-unit 
quantities, the terms impedance and impedance voltage 
are interchangeable. 



310 



The vector sum of the leakage reactance and the 
resistance of the windings comprise the impedance of a 
transformer. Figure 12.11 shows that the X/R ratio of a 
typical transformer increases as the size (megavoltampere 
rating) increases (13). The transformer impedance is ex- 
tremely important since it plays a major role in voltage 
regulation as well as dictating the amount of current that 
can be delivered to a fault on the secondary. Here, the 
maximum symmetrical secondary-fault current (three- 
phase fault) can quickly be estimated by 



L, 



u 



J Tpu 



(12.2) 



where I flmax) = maximum symmetrical fault current, A, 
I FLs = full-load current of transformer secondary, 
A, 
and Z.^ = transformer per-unit impedance, pu fl. 

For the previous example, 



6 -. 

I— XIK 

UJ — ' 
<-J O 

UJ 



40 
30 
20 
10 







































































































ty. 



) 1 2 5 10 20 50 100 200 500 1,000 

SELF -COOLED TRANSFORMER RATING, MVA 

Figure 12.11.— Typical X/R ratio versus transformer capaci- 



ty \ 'A 
V3kV 



500 



V3 (0.48) 



= 601 



and 



K 



601 
0.06 



= 10,023 A. 



This technique produces a more pessimistic estimate than 
would actually occur, since it assumes that the trans- 
former primary is connected to an infinite bus. A more 
realistic value for fault current should take into account 
the system impedance upstream from the transformer as 
well as an equivalent impedance for the utility. However, 
this system impedance is constantly changing because of 
the dynamic nature of the mining process. By using the 
above method, it is possible to obtain a conservative 
determination of the required interrupting capacity for 
protective circuitry serving the utilization equipment. 

For transformers used in mine power centers, the 
per-unit values of impedance normally lie in the following 
range: 

• 0.04 to 0.05 pu for 150 to 450 kVA, and 

• 0.05 to 0.06 pu for 500 to 1,250 kVA. 

The 0.04-pu lower limit is usually needed to limit short- 
circuit currents to reasonably safe values. For large power 
transformers, the leakage reactance may be intentionally 
increased in the design of the transformer to further limit 
the available fault current. External air-core reactors in 
series with the secondary windings are sometimes used to 
restrict fault current from transformers that feed rectifi- 
ers. This will be discussed in the section on ac-dc combi- 
nation power centers. 

For transformers used in mine power centers, the 
following minimum insulation ratings should be used (the 
BIL's in parentheses are optional): 

4.16-kV primary, 35-kV (60-kV) BIL; 
7.2-kV, 60-kV (75-kV) BIL; 
12.47-kV (13.2 and 13.8 kV), 95-kV BIL; and 
23-kV, 150-kV BIL. 

Coordination with surge arrester characteristics must be 
maintained, as discussed earlier. (See chapter 11 for fur- 
ther information.) 



The kilovoltampere rating of a transformer is based 
upon continuous operation at rated voltage and frequency 
without exceeding the specified temperature rise or limit- 
ing temperature. The temperature rise depends upon the 
core and conductor losses. The core losses remain constant 
with load, while the conductor losses are determined by 
the I 2 R of the windings. 

Dry transformers used in mine power centers should 
have class 220 °C insulation. This classification terminol- 
ogy has replaced the old class H 150°C rise notation that 
was based on an average ambient temperature of 30°C, 
allowing a maximum 150°C temperature rise in the 
windings (2). This restricted the average insulation tem- 
perature to 180°C, but a 220°C hot-spot temperature was 
also allowed. The new classification simply places an 
absolute allowable maximum temperature on the trans- 
former when it is operating continuously under full capac- 
ity at rated voltage and current. 

Because of the use of trailing cables and mobile equip- 
ment, short-circuit conditions are frequently encountered in 
mining, and transformer windings can undergo considerable 
thermal and magnetic stress during these occurrences. Con- 
sequently, the transformer should be designed to withstand 
25 times rated current in any winding for a period of 2 s. One 
method of curtailing expansion is to add an insulated brace 
across the windings, and some engineers request adjustable 
braces, which allow periodic readjustments to be made when 
expansion causes loosening. 

A temperature-sensing device (fig. 12.2, No. 10) can be 
placed in the transformer windings to prevent damage 
from overheating. The device controls a set of contacts 
located in the pilot circuit of the incoming distribution 
line. If the transformer temperature exceeds the specified 
limit, the contacts in the pilot circuit are opened, which 
results in tripping the upstream circuit breaker. As an 
alternative, the contacts can be used to activate the 
tripping element of a main secondary breaker or all 
outgoing breakers. 

Mine power centers are usually cooled by natural 
convection, and the side panels of the transformer com- 
partment normally have louvers to allow air circulation. 
The transformer windings are designed so that the heat 
generated by the I 2 R losses is exposed to an adequate 
amount of cooling to handle the expected loads. The 
effective cooling areas are the inside of the winding, the 



311 



outside of the winding, and the cooling ducts within the 
winding. The movement of air by convection carries away 
the heat generated by the winding (13). 

Transformer Construction 

The two main transformer components are the core 
and the windings. Figures 12.12 and 12.13 show a typical 
mine power-center transformer in the process of being 
constructed and in completed form. The transformer core 
provides a path of low reluctance to the flux produced by 
the primary windings. In effect, it comprises the magnetic 
circuit of the transformer. To reduce eddy currents, the 
core is constructed of laminated sheet steel. The eddy- 
current losses vary as the square of the thickness of 
laminations, and core laminations are usually from 0.01 to 
0.02 in thick (13). The laminated material is cut from 
silicon-iron sheets. The silicon reduces the reluctance to 
hysteresis and prevents increased loss with age. The 
laminated material is specially annealed to obtain a high 
permeability and is also treated with a chemical coating to 
insulate the laminations from each other. The laminations 
are stacked one upon another to construct a closed mag- 
netic path. Alternate layers are staggered in an interlam- 
inar core-and-gap construction so all joints do not meet at 
the same place. This is done to reduce the effect of the air 
gap between the joints and make the entire structure 
function more like a solid piece of iron (13). 

The primary and secondary windings comprise the 
current circuit of the transformer. The windings are de- 
signed to get the required number of turns into the 
minimum of space through the core opening, while also 
providing enough room for insulation and cooling ducts. 
Transformer windings are made of copper or aluminum. 
Because of its lower conductivity, aluminum requires a 
larger cross section of conductor and thus a larger opening 
in the core. The conductors used in the windings may be 
round, square, or rectangular in cross section. Aluminum 
secondaries are often wound using sheet metal. All wind- 
ings may be insulated with enamel, mica paper, NOMEX 
paper, silicon glass tape, or a combination of these materi- 
als (8). 

Faraday Shields 

Transformer windings can be provided with a 
grounded Faraday (or electrostatic) shield to destroy inter- 
winding capacitance between the primary and secondary, 
thus reducing the danger of transferring distribution 
transients to utilization. Another important use of the 
shield is to prevent interwinding faults between layered 
windings. This is especially critical when the secondaries 
are isolated above ground, as the primary voltage can be 
impressed on the secondary without detection. Using the 
shield, a ground fault will occur in this situation, and the 
resulting ground current will be detected by upstream 
ground-fault relaying. 

The shield consists of a layer of nonmagnetic metal 
placed between the primary and secondary windings, 
insulated from all windings, and connected solidly to 
ground (1). Aluminum or copper can be used, but the 
shield should be made of the same material as the main 
windings. Physically, it may be a single turn of sheet 
metal or a closely wound single layer of wire. To obtain 
interwinding fault protection, the shield must have the 
same ampacity as the grounding conductors leading to the 




Figure 12.12.— Typical mine power-center transformer under 
construction. (Courtesy PEMCO Corp.) 




Figure 12.13.— Completed transformer prior to installation. 
(Courtesy PEMCO Corp.) 



power center, in order to carry maximum available 
ground-fault current (at least, equivalent to one-half the 
cross-sectional area of the primary-winding wire). 



GROUNDING RESISTOR 

Federal regulations require the maximum frame po- 
tential on medium-voltage and low-voltage circuits to be 
limited to 40 V. In an attempt to ensure this, the regula- 
tions further require the maximum ground-fault current 
of these circuits to be limited to 25 A. However, many 
modern power centers are designed to limit the ground 
fault current to 15 A for an additional factor of safety. 
Limiting the ground-fault current results in the following 
additional benefits (7): 

• Reduction in burning and melting of faulted elec- 
trical equipment, 

• Reduction in mechanical stress of faulted electrical 
equipment, 

• Ability to selectively clear the faulted circuit, and 

• Reduction in overvoltages that might cause insula- 
tion failure. 



312 



The grounding resistor (fig. 12.2, No. 11) is inserted 
between the neutral of the transformer and the power- 
center frame as a means of limiting the ground-fault 
current. The ohmic value of the grounding resistor is 
based on the maximum ground-fault current condition; 
that is, a ground fault at the secondary terminal of the 
transformer. For this situation and neglecting the trans- 
former impedance, the resistance is determined by 



R V >n 



(12.3) 



where R G = ohmic value of grounding resistor, U, 

V ln = line-to-neutral potential of transformer sec- 
ondary, V, 
and If = maximum ground-fault current, A. 

As an example, consider sizing a grounding resistor to 
limit the ground-fault current to 15 A for a 480-V system. 
The ohmic value of the resistor would be 



Rn = 



480/V3 
15 



= 18.5 n. 



When sizing a grounding resistor, the time rating 
must also be considered. Under normal conditions, an 
insignificant amount of current flows through the resistor, 
but during the worst case situation, 15 A may flow 
through it. Thus the power (P) dissipated by the neutral- 
grounding resistor would be 



or 



P = I f 2 R G 
P = (15) 2 (18.5) = 4,162 W. 



(12.4) 



Normally the resistor would only be required to dissipate 
this power for the time it takes for a circuit breaker to trip, 
but the possibility of the circuit breaker's failing to trip 
must also be considered. Federal regulations require the 
resistor rating to be based on an extended time rating. 
This has been defined as 90 days of operation per year. The 
end of the resistor that is connected to the neutral of the 
transformer is also required to be insulated for line-to-line 
voltage. 

BUSWAY 

The power-center output requires numerous taps for 
feeding the utilization circuits. A busway (or busbars) 
provides a convenient and economical means of providing 
these taps (fig. 12.2, No. 12). The busway consists of flat, 
bare conductors supported within the power-center enclo- 
sure by means of insulators made of glass, polyester, or 
porcelain, as can be seen in figure 12.14. Busways are 
available with either copper or aluminum conductors. 
Aluminum has a lower electrical conductivity and lower 
mechanical strength and quickly forms an insulating film 
on its surface when exposed to the atmosphere. If alumi- 
num conductors are used, they should have electroplated 
contact surfaces (tin or silver), and bolting practices that 
accommodate aluminum mechanical properties should be 
used at electrical joints (5). 

The continuous-current rating of the busway is based 
on the cross-sectional area of the conductor and a maxi- 
mum temperature rise of 55° C from ambient. The short- 




Figure 12.14.— Typical bus work in power center under con- 
struction. (Courtesy PEMCO Corp.) 



circuit rating should exceed the maximum available fault 
current. The sizing of the busway is basically the same as 
that discussed for other conductors in chapter 8. 



OUTGOING CIRCUIT BREAKER 

Molded-case circuit breakers (fig. 12.2, No. 13) are 
usually employed to protect ac utilization equipment and 
associated cables. Some manufacturers have special mine- 
duty breakers available, which have greater ruggedness 
and reliability to meet the demands of mining practices. 
Dual-element fuses are rarely used because of single- 
phasing problems and the lack of auxiliary protective- 
relaying capabilities. A complete discussion of molded- 
case breaker and fuse applications, sizing, and problems is 
available in chapters 9 and 10. A presentation of molded- 
case devices with solid-state tripping elements is con- 
tained in chapter 14. 

The power-center circuit breaker compartment must 
be designed for easy access, but protection is necessary 
against accidental exposure to energized terminals and 
conductors. Dead-front panels are the preferred method of 
construction. Here the breaker is mounted on a recessed 
panel so that only the operating handle and adjustments 
are accessible from the outside. As an alternative, the 
breakers can be surface mounted, but here too a barrier 
should cover the power terminals. Doors should be used on 
the exterior frame to minimize exposure of the compart- 
ment to dust and other contaminants. 

The two kinds of protection that can be provided 
directly by molded-case circuit breakers are short circuit 
and overload. Short-circuit protection is required on all 
outgoing power (ungrounded) conductors, and maximum 
allowable instantaneous-trip settings are established by 
Federal regulations. Overload protection is mandated in 
some States. Magnetic elements give instantaneous-trip 
protection, and thermal elements afford overload protec- 
tion. Pickup settings for both are based on the smallest 
size of cable being protected. 

Undervoltage protection of each outgoing circuit is 
almost always required, the only exception being if the 
protected equipment has its own. An undervoltage release 
(UVR) is invariably used and is an auxiliary solenoid that 
trips the operating mechanism of the breaker whenever its 



313 



coil voltage drops below 40% to 60% of rated. Additional 
external protective relaying can be given to the outgoing 
circuit through the UVR. Note that even though under- 
voltage protection may be needed, each outgoing breaker 
in a mine power center should contain an UVR. Outgoing 
breakers are rarely provided with shunt-trip solenoids. 

Terminal connectors are available for molded-case 
breakers for either single- or multiple-conductor entry. 
Copper conductors require the use of copper terminals, 
and aluminum conductors require aluminum-compatible 
terminals. If multiple-conductor terminals are employed, 
the terminal must be constructed so that each conductor 
can be tightened without removing another conductor. 
Figure 12.15 is a representative view of conductor connec- 
tions to a molded-case breaker. 

The selection of molded-case circuit breakers is based 
on the voltage, frequency, interrupting capacity, 
continuous-current rating, and trip settings. Even though 
this topic is presented in chapters 9 and 10, additional 
coverage of tbe practical considerations for interrupting- 
capacity and continuous-current selection is warranted 
here. 

Molded-case circuit breakers often begin interrupting 
short-circuit currents during the first cycle after the fault, 
and so they must be selected on the basis of maximum 
first-cycle asymmetrical fault current. The breakers are 
usually rated on a symmetrical current basis, which 
eliminates applying dc offset multipliers when selecting 
the breaker. It is generally considered adequate to use 
calculated symmetrical short-circuit currents for load- 
center applications. 

As an example, consider a 750-kVA power center with 
a secondary utilization voltage of 600 V. If the transformer 
impedance is 6%, the maximum symmetrical fault cur- 
rent, If (max) , can be calculated as 

750 
1fLs " V3 (0.6) " 722 A ' 



722 
■<"■"■■«> = O06- 12 ' 033A - 



This of course assumes a worst case condition, since 
transformer impedance is the only limiting factor for all 
fault current. In this case, even a 100-A standard mine- 
duty breaker could interrupt the fault; the typical inter- 
rupting capacity is 14,000 A symmetrical at 600 V. How- 
ever, if the power center contains a 1,000-kVA transformer 
with 6% impedance, fault current becomes 



*- = ^S = 962 a > 



962 
0.06 



= 16,038 A. 



Under this situation, the standard 100-A molded-case 
breaker would have inadequate interrupting capacity. (A 
premium-duty 100-A unit would be safe; typical interrupt- 
ing capacity is 18,000 A symmetrical at 600 V. See chapter 
9 for other typical ratings.) The simple example given here 
shows just one reason for the minimum transformer im- 
pedance stated earlier in this chapter. 




Figure 12.15.— Typical conductor connection to molded- 
case circuit breaker. (Courtesy PEMCO Corp.) 



The continuous-current rating for a molded-case cir- 
cuit breaker is actually the rating of the thermal-trip 
element, which can be less than the breaker frame size (a 
thermal-magnetic breaker). Magnetic-only devices have a 
continuous-current rating equal to the frame size. The 
rating is based upon 100% continuous current at 40°C, but 
sizing the breaker is normally related to 80% of the rating. 
This means that sizing is based upon 1.25 times the cable 
ampacity or the full-load circuit current. In other words, the 
breaker operates at 80% of the continuous-current rating at 
full-load current. (Note that some molded-case breakers are 
applied at 100% full-load current, particularly those with 
solid-state tripping elements.) 

For instance, consider sizing a breaker to protect a 4/0 
cable for overload. A typical ampacity for 4/0 three- 
conductor unshielded trailing cable is 287 A. Thus, the 
breaker continuous-current rating would be (1.25) (287) or 
359 A. The next highest standard size, or a 400-A rating, 
would then be selected. However, when sizing outgoing 
breakers for a mine power center, the method of selection 
is not so simple. The process can approach something of an 
art based on prior experience, especially when under- 
ground coal mining equipment is involved. An actual 
situation is used here to illustrate some difficulties that 
can be encountered. 

From a practical standpoint, 4/0 is the largest 
trailing-cable size that can be used in underground mining 
and is the most common conductor size for continuous 
miner applications. Assume that the continuous miner 
has the following motors; 

• Cutting motors, two at 90 hp; 

• Pump motor, 50 hp; 

• Gathering-head motor, 70 hp; and 

• Traction motors, two at 35 hp. 

Assuming a 480-V system, an average efficiency of 90%, 
and an average power factor of 90%, the maximum theo- 
retical full-load current for the machine (370 connected 
horsepower) would be from equation 8.3: 



I™ = 



(370) (0.746) 
V3 (0.48) (0.9) (0.9) 



= 410 A. 



This current exceeds the cable ampacity; thus, breaker 
sizing based on cable ampacity is not applicable. It should 
be realized that it would be a rare situation to have all 



314 



motors draw rated current. Each motor or motor pair has 
its own noncontinuous duty cycle, which is a function of 
many variables. It is for this reason that an undersized 
trailing cable can be used. Yet the circuit breaker must be 
selected to handle the load demands of the continuous 
miner. Therefore, the sizing of the continuous-current 
rating for a breaker does not follow a set procedure. Most 
power-center manufacturers are well aware of the machin- 
ery available and can usually provide the optimum circuit- 
breaker size for specifications. (Again, see chapter 9 for 
more discussion.) 

A main circuit breaker (fig. 12.2, No. 14) is often 
recommended if the number of outgoing circuits exceeds 
three. The breaker has the bus work as its primary 
protection zone, but also serves as a backup to the outgo- 
ing breakers. The rating can be based upon the full-load 
current of the transformer secondary or the ampacity of 
the bus work, whichever is lower. In general, the bus work 
is sized on the transformer full-load current. For example, 
the full-load current for the 600-V secondary of a 750-kVA 
power center is 



750 
V3 (0.6) 



= 722 A. 



Thus, the breaker continuous-current rating should be 
based on 

I FL = (1.25) (722) = 902 A. 



A 1,200- A frame with a 1,000- A continuous-current rating 
would be selected. The thermal elements would provide 
overload protection to the transformer and bus work; 
magnetic elements could be set to give short-circuit pro- 
tection to the bus work. Some main breakers are provided 
only with magnetic elements. Obviously, the main breaker 
must be coordinated with the outgoing breakers and the 
high-voltage fuses on the transformer primary. 



schemes work with a low-cost, low-burden CT. Since the 
CT is drastically underutilized, its secondary current 
cannot be predicted by knowing the turns-ratio of the CT. 
The pickup values must be determined by testing. 

The normally open (NO) contacts of the relay usually 
parallel the UVR of the associated circuit breaker. When 
the relay pickup value is exceeded, its associated contacts 
close and short out the UVR coil. This technique has been 
adopted to eliminate nuisance tripping due to bounce and 
vibration, which hamper circuits that have contacts in 
series. One problem with paralleled contacts as shown in 
figure 12.16 is that ground-fault protection is lost if the 
relay is removed from its socket. To prevent this, UVR 
power can be supplied through a jumper in the ground- 
fault relay case (fig. 12.17). Without the relay, the circuit 
breaker cannot be closed, except in some small molded- 
case units, such as 50-A units. 



Line conductors 
Molded-case circuit breaker 




To control 
transformer 



GTR(50G) 



Figure 12.16.— Zero-sequence relaying on outgoing circuit 
with control connections to breaker. 



GROUND-FAULT PROTECTION 



As mentioned earlier, a grounding resistor is placed 
between the transformer neutral and the power-center 
frame to limit ground-fault current to not more than 25 A. 
Ground-fault relaying must be used on each outgoing ac 
circuit to initiate circuit interruption during malfunc- 
tions. The common relay schemes for ac applications in 
load centers are zero-sequence, neutral (direct), and poten- 
tial. Zero-sequence relaying is the most common ground- 
fault protection and is utilized on practically all outgoing ac 
circuits. As shown in figure 12.16, the three line conductors 
are passed through a window-type current transformer (CT), 
and burden for the CT secondary is a ground-trip relay. Relay 
operation was described in chapter 9. 

The ground-trip relay must be set to pick up at 40% or 
less of the maximum ground-fault current. For this situa- 
tion, a low-ratio CT might be expected to be used, such as 
25:5 or 50:5 (ampere-turns ratio). However, better sensitiv- 
ity has been realized by using a high-ratio CT, 350:5 or 
higher, with a voltage-sensitive relay (about 1.5-V pickup), 
instead of a current-sensitive relay for the tripping device. 
Some manufacturers employ a slight variation to this 
conventional scheme by rectifying the CT output and 
using a dc voltage-sensitive relay. Either way, these 



Line conductors 

^- Molded-case circuit breaker 




To control 
transformer 



GTR(50G) 



Figure 12.17.— Zero-sequence relaying with jumper in relay 
case. 



315 



Neutral relaying is sometimes used in mine power 
centers when a main breaker is used. The neutral conduc- 
tor from the transformer is encircled by a CT, as shown in 
figure 12.18. The CT and the ground-trip relay are the 
same as discussed for zero-sequence relaying. However, a 
time delay should be introduced by the ground-trip relay 
so selective interruption of the faulted circuit occurs. If the 
circuit breaker of the faulted circuit fails to trip, the main 
circuit breaker will provide backup protection after the 
prescribed time delay. The time delay is on the order of 0.5 
s and can be achieved by pneumatic or electronic means. 
The breaker trip device may be either a shunt trip or UVR, 
but the UVR is preferred. 

Potential relaying can also be used as backup protec- 
tion in conjunction with a main circuit breaker. Unlike 
zero-sequence and neutral relaying, potential relaying has 
the advantage of being able to detect a ground fault with 
the neutral grounding resistor open. Figure 12.19A shows 
a potential-relaying scheme using a PT for obtaining a 
voltage within the operating range of the relay whereas 
figure 12.19B shows a voltage divider to provide the same 
function. 

For both schemes, the maximum pickup voltage of the 
relay should correspond to the voltage developed across the 
relay at 40% of the maximum fault current. When using a 
voltage divider circuit, care must be taken in sizing the 
resistors so the total equivalent resistance of the parallel 
combination is not reduced significantly. The combined 
value (R x + R 2 ) must be significantly higher than the 
value of the neutral grounding resistor. The impedance of 
the relay should also be significantly higher than the 
resistance of R 2 for the same reason. The power (I 2 R) 
rating of the resistors should be capable of withstanding 
the currents expected through them under a maximum 
ground-fault condition. 

Figure 12.19C illustrates a popular alternative to 
backup relaying. Here the relay coil is replaced by a red 
warning light located in a rugged or explosion-proof hous- 
ing mounted on the outside of the power center. The light 




I 



r r r 
III 



UVR 



•— Wv *• 



To control 
transformer 



Main circuit breaker 



Figure 12.18.— Neutral relaying applied to grounding- 
resistor current as backup protection. 



bulb voltage is matched to the PT to give maximum 
brilliance under maximum ground-fault conditions but 
noticeable light output at 40% of current limit. Many 
engineers prefer this technique as it provides an obvious 
indication of ground fault and is especially useful when 
primary ground-fault relaying is inoperative either 
through tampering or malfunction. 

Test circuits for ensuring pickup at 40% of the maxi- 
mum ground-fault current can be incorporated into the 
power center: figure 12.20 is an example. A control trans- 
former with a secondary voltage of 12 V is used as a 



Neutral 

grounding 

resistor 




To circuit 
breaker trip 



A Potential transformer 



Neutral 

grounding 

resistor 




_L To circuit 

nr breaker trip 



B Resistive voltage divider 



Neutral 

grounding 

resistor 




-Ground-fault 
indicator 



C Warning light 

Figure 12.19.— Backup protection devices associated with 
mine power centers. 



Line conductors 
CT 



To control 
transformer 




1 To circuit 
breaker trip 



120/12 V 
50 VA 



6 turns 



GTR 



Figure 12.20.— Typical test circuit for zero-sequence relay- 



ing. 



316 



current source. The 12-12 resistor is inserted in the circuit 
to limit the secondary current to 1.0 A. The six turns 
through the CT at 1.0 A produce the same effect as 6.0 A 
of zero-sequence current for the power conductors. The 
current corresponds to 40% of a 15-A system so that 
depressing the test button simulates a ground fault at 40% 
of the maximum current. The associated circuit breaker 
should then trip. 



SINGLE-PHASE TRANSFORMERS 



■« — 


-EZD ;> 


To 
480 V 


K 


-* 


-n n — -> 



240/480 V 

50 VA 

120 V 




To 
ground 
monitor 



Figure 12.21.— Simple control circuit incorporating one 
ground-fault relay (GTR) and one ground-check relay (GCR). 



Single-phase transformers are used in power centers 
to supply 120 V to the control circuit and 240/120 V to 
convenience outlets. The control circuit consists of ground- 
monitoring systems, undervoltage releases, and ground- 
fault circuitry for each associated machine circuit, as well 
as relay connections for other protection devices. Conve- 
nience outlets can be used for portable power tools, area 
lighting, or external test instrumentation. Control trans- 
former capacity normally ranges from 50 to 500 VA, 
whereas transformer capacity for 120/240-V outlets often 
falls in the range of 5 to 10 kVA. Larger capacities may be 
found, particularly in mine power centers with extensive 
control circuitry. A simple control circuit and a 
convenience-outlet circuit are shown in figures 12.21 and 
12.22, respectively; in both circuits, fuses are used to 
protect the transformer primaries. Circuit breakers can be 
employed to protect convenience-transformer primaries, 
but fuses are recommended for control-power circuits be- 
cause breakers can be tripped by an unwary miner, thus 
deactivating the load center. 

Control-power fuses can be mounted in insulated 
dead-front holders or in a typical spring-clip arrangement 
that is only accessible to authorized personnel through a 
bolted cover. These two mounting types are illustrated in 
figure 12.23. The testing and changing of control-circuit 
fuses are among the more frequent electrical maintenance 
procedures for all types of equipment. Typical spring-clip 
fuse holders have uninsulated exposed metal clips, and 
rushed repair persons frequently remove the fuses from 
these holders for testing without deenergizing the circuits, 
thus placing their hands within close proximity of the 
energized clips. A slight inadvertent movement of the 
hand can result in electric shock. With dead-front fuse 
mounting, however, all energized components are enclosed 
in an insulated housing so that the fuses can be removed 
and replaced without exposing the electrician to the 
metallic clips and fuse ends. 

A variety of schemes exist for handling the control- 
and utility-circuit voltages. Each machine circuit can have 
its own individual control transformer, but it is more 
common for a single control transformer to supply all the 
control circuits. When all control circuits are supplied by a 
single source, the control-circuit transformer can be elim- 
inated, with the control voltage supplied by the secondary 
of the convenience-outlet transformer. Many mine power 
engineers dislike this combination because a failure or 
misuse of a convenience-outlet circuit could result in 
blowing the transformer fusing. This in turn would cause 
the loss of control power and all power to the mining 
machinery would stop. For this reason, some engineers 
prefer not to incorporate any 120/240-V outlets in their 
load centers, even with separate transformers. 




240-V receptacle 
120-V receptacle 



480V-120/240-V 
5kVA 



15A 



Figure 12.22.— Simple convenience-outlet circuit for 120- or 
240-V single phase. 



Dead-front 
panel 




X 



Insulated 
fuse housing 



Typical mounting Dead-front mounting 

Figure 12.23.— Fuse mountings. 

METERING CIRCUITS 

The mine power center should include metering cir- 
cuits to monitor line voltages and currents of all three 
phases (fig. 12.2, Nos. 15-16). The built-in instrumenta- 
tion is an invaluable aid to maintenance, as it allows a 
firsthand look at the electrical operation of the load center 
and gives a composite view of the loads it serves. The 
metering is usually for distribution or utilization voltages 
and current, but rarely for both. 

Figure 12.24 illustrates a common approach for me- 
tering voltage, where potential transformers (PT's) are 
used to reduce the line voltage to a value within the rating 
of the voltmeter potential coil. Two single-phase trans- 
formers are connected in open delta, and a four-position 
switch allows the three line-to-line voltages to be moni- 
tored in addition to the o/f position. 

As shown in figure 12.25, two CT's are sufficient for 
metering the current in all three lines and isolating the 
meter circuit from the line. The CT ratio should be as low 



317 



600 -V bus 




Voltmeter 
switch 



^Hf 



V 0-750 V 



__^2_3_4 
g-b X X 
b-c X X 
c-a X X 



750:150 
25 VA 



Figure 12.24.— Typical metering circuit for line-to-line 
voltages. 



■H| — LZZr 





1 


2 


3 


4 


- 


X 






X 


a 


X 






X 


- 


X 


X 




X 


b 




X 




X 


- 




X 


X 




c 




X 


X 




- 


X 


X 


X 




OFF 


X 




X 





1,500:5 




a b c "=■ 

Figure 12.25.— Typical metering circuit for line currents. 



as possible, without exceeding rated current in the second- 
ary winding. A common recommendation is that the ratio 
be such that normal CT secondary current is 0.5 to 0.75 of 
the full-scale rating of the meter (5). The CT turns ratio 
can cause dangerously high voltage on the secondary if it 
is opened, so the ammeter switch must be designed so the 
secondary is not open circuited during the transition from 
one switch position to another. These are known as short- 
ing switch contacts. 



OUTGOING CABLE COUPLERS 

The trailing cables for utilization equipment are al- 
most always connected to the power center by means of 
low-voltage or medium-voltage cable couplers (fig. 12.2, 
No. 17). The couplers are rated at 600 V for low- voltage 
operation and at 1,000 V for medium voltages. Standard 
current ratings are 100, 225, 400, and 800 A, but larger 
sizes are readily available. Additional information can be 
found in the coupler section of chapter 8. 



The outgoing cable couplers on the power center are 
female-contact receptacles, which connect with the male- 
contact plugs of the trailing cables. When disconnecting, 
the opening of contacts must follow a sequential order. The 
couplers serve as a visual disconnect for the low-voltage 
and medium-voltage circuits served by the mine power 
center, and for safety the couplers should always be locked 
out when not in use. Lockout is the process of deenergizing 
a circuit so that it cannot be energized without authority. 
Improper lockout or failure to lock out has been a leading 
cause of electrical accidents. For example, a repair person 
deenergizes a faulty circuit at the power center prior to 
performing repair work. While the repair person works on 
the faulty circuit, another worker who is unaware of the 
situation mistakes one cable for another, energizes the 
faulty circuit, and subjects the repair person to electric 
shock or electrocution. 

Effective lockout can be provided by routine use of 
locking dust covers or keyed couplers. With keyed couplers, 
the receptacle of each outgoing circuit is matched to fit 
only one cable-mounted plug so that mistakes cannot be 
made. On the power center, the dust covers are hinged to 
the receptacle or connected by a chain. The hinged covers 
are preferred from a safety standpoint since they are less 
likely to become disengaged and lost. Locking dust covers 
on cable-mounted plugs offer even better protection, as 
once the plug is locked it is impossible to connect it to any 
receptacle. The covers also prevent damage and dirt accu- 
mulation when the plugs are not in use. Although chain- 
connected locking dust covers are the only type presently 
available for plugs, hinged lids would again offer improved 
safety. 

Another lockout technique is to drill a hole in the 
spare pilot-circuit pin of the cable-mounted plug. A pad- 
lock can then be inserted in the hole by the electrician. 
This technique is extremely simple and inexpensive, but it 
provides excellent lockout protection by prohibiting plug- 
receptacle coupling. 

The installation of two receptacles for each outgoing 
machine circuit can also provide lockout protection. One 
receptacle is used for powering the outgoing circuit while 
the other is used strictly for locking out the circuit when 
necessary. Actually, the lockout receptacle is not a true 
receptacle since its housing need not contain internal 
contacts. The lockout receptacle can also protect the plug 
from damage while the associated circuit is not in use. 

GROUND-CHECK MONITORS 

Low-voltage and medium-voltage resistance-grounded 
systems are required to have a fail-safe ground-check 
circuit to continuously monitor the continuity of the 
grounding conductor. The monitor must cause its associ- 
ated circuit breaker to trip if the grounding conductor or a 
pilot wire is broken. An indicator lamp on the monitor 
should indicate a tripped condition. The monitors are 
usually enclosed in a dead-front package and mounted 
near the associated circuit breaker. As was seen in chapter 
9, monitors in common use in mining can be divided into 
two general classifications: impedance types and continu- 
ity types. 

Impedance monitors require the trailing cable to have 
a pilot conductor. The monitor is calibrated to the imped- 
ance of the loop formed by the pilot and grounding 
conductors, and the device then monitors the change of 
impedance from the initial calibration. If the impedance of 



318 



the loop increases beyond a preset value, the monitor must 
trip its associated circuit breaker by opening a set of 
contacts in series with the undervoltage release. 

The maximum allowable increase in impedance is 
dependent upon the maximum ground-fault current per- 
mitted by the system. For a 15-A neutral-grounding resis- 
tor, the monitor should cause tripping if the impedance 
increases by 2.7 ft. This value is based upon the maximum 
allowable frame-to-ground potential of 40 V, as follows: 



40 



Z = -f* m ■ 



From the above, it is apparent that a ground-check moni- 
tor does not ensure that the frame potential of a piece of 
equipment will not rise above 40 V, since the device 
monitors only the change in impedance of the pilot and 
grounding-conductor loop and not the actual impedance of 
the grounding conductor. 

A schematic of a common impedance monitor is 
shown in figure 12.26. This particular monitor is powered 
from a 24- to 32-V source, but others use 120-V power. 
Out-of-phase induced currents during motor starting can 
result in cancelling out the monitoring current, which in 
turn can cause nuisance tripping. As a result, a polarity- 
reversal switch should be provided to change the phase 
relationship of the pilot current with respect to the in- 
duced current. Some manufacturers use impedance- 
matching transformers to amplify the change in imped- 
ance for easy detection. The monitor should also provide a 
test button. With the button depressed, the appropriately 
sized resistor is inserted into the pilot circuit, which 
should result in tripping the circuit breaker. A disadvan- 
tage of the impedance monitor is that it cannot detect a 
pilot-to-ground fault. This type of monitor is also suscep- 
tible to problems with parallel paths. 

Continuity monitors do not monitor impedance change, 
but only the grounding-conductor continuity. However, they 
must be adequately immune to parallel paths. Continuity 
motors are audio units and do not require a pilot conductor 
for operation. Figure 12.27 contains a block diagram of a 



common unit. Most makes can also be wired for pilot 
operation (fig. 12.28), but only the operation of the pilotless 
configuration will be discussed. 

The monitor generates an audio frequency, which is 
coupled to the grounding conductor by means of the 
transmitting coil. The pilot wire is eliminated by using the 
line conductors as a return path. Filters at the monitoring 
location and within the monitored machine are necessary 
for coupling and uncoupling the audio signal from the line 
conductors. If the grounding conductor is intact, the re- 
ceiver coil picks up the audio signal. If the grounding 
conductor is open, the receiver coil will pick up no signal, 
and the monitor will cause a set of contacts to open the 
undervoltage circuit of its associated circuit breaker. 
Chapter 9 contains a discussion of a continuity monitor 
that operates in a somewhat different fashion; the signal is 



r 



i/i/hmi. 

Frame 
ground 



■g Large 
coupler 



I" 



Line 



Line 2 



Transmitter 
coil 



Line 3 



Frame 
ground 



-PP£ 



=>PPr 



Grounding 
conductor 



Small 
coupler 



"1 



Frame 
ground 



I3J 



Receiver 
coil 



MONITORED 
MACHINE 



Ground 
monitor 



To breaker 
trip circuit 



To control | 
power(l20Vac)l 



r_j 



POWER-CENTER DISTRIBUTION 
BOX OR OTHER 

Figure 12.27.— Block diagram of continuity monitor con- 
nected in pilotless mode. 



32 Vac 



Polarity switch 




Transformer 



Al LI 
Deenergized 



3.0-11 resistor B 



0@ 

Figure 12.26.— Typical impedance monitor circuit. 



r 



Transmitter coil 

rrn 



~l 



Frame 
ground 



Pilot 
conductor 



Receiver 
coil 



Frame 
ground 



=PP>- 



Groundmg 
conductor 



Frame 
ground 



Ground 
monitor 



I 



To breaker 
trip circuit 



I I 

MONITORED 
MACHINE 



To control 
power (120 Vac) 



SUBSTATION, POWER-CENTER 
DISTRIBUTION BOX, OR OTHER 



Figure 12.28.— Block diagram of continuity monitor wired 
for pilot operation. 



319 



impressed on and removed from the line conductors, with 
the grounding conductor used for the return path. 

The major advantage of the continuity monitor is that 
it is immune to parallel paths and stray currents. It is also 
immune to pilot-to-ground faults if a pilot conductor is not 
used. However, the continuity monitor is expensive and 
complex when compared with an impedance monitor. A 
pilotless monitor must still be wired into the pilot and 
grounding contacts of the coupler in order to trip the 
circuit breakers when disconnecting the coupler. Another 
problem is that the grounding conductor must be isolated 
from the coupler shell, or the intended monitoring is 
bypassed. For metallic shells, a separate grounding con- 
ductor must be supplied through a spare coupler contact. 



POWER-FACTOR CORRECTION 

Some mining machines have notoriously poor power 
factors resulting from underutilization of induction mo- 
tors. Perhaps the most outstanding example is the contin- 
uous miner, which can have a power factor that averages 
0.6 lagging during the operational cycle. Whether it is this 
machine or others that create excessive reactive power, the 
result is poor power-system efficiency and utilization. If 
the power factor is poor at the purchase points (under 0.80, 
for example), the utility company will attach a penalty to 
the power bill. 

There are many opinions on the best location at which 
to improve power factors. The most obvious place is at the 
utilization points. This strategy is common in surface 
mines, particularly with excavating machines. However, 
surface mining equipment has adequate space to incorpo- 
rate the correction equipment, an advantage not common 
in underground mining machines. If the improvement is 
made at a convenient location immediately upstream from 
offending motors, the entire system up to and including 
the substation can benefit through lower apparent power, 
reduced line currents, and better voltage regulation. This 
location is commonly on the utilization buswork in mine 
power centers. On the other hand, if correction is not 
attempted here, power factors will be increased by the 
inherent distribution-system capacitance (for example, 
that associated with shielded cables) or somewhat can- 
celled by the diversity of operation that always exists 
between mining machinery and equipment sections. As a 
result, some engineers would rather correct the overall 
power factor at the distribution side of the substation. 

Consequently, three general locations can be consid- 
ered if power factors need correction: the machine, in the 
power centers, or at the substation. Each individual mine 
power system must be analyzed to find the solution, and 
the final decision must consider both electrical operation 
and economics. One consideration should be the cost of 
load-center correction versus correction at the substation 
versus the penalty added to the power bill. If a decision is 
made to add correction, the common approach is to use a 
bank of capacitors, also known as a static bank. (Synchro- 
nous rotating machinery is not logistically practical for 
most cases.) 

Precautions must be taken when applying correction- 
capacitor banks, regardless of the location. They are a 
system load, and thus the bank must not be grounded on 
resistance-grounded systems. In the past, the common 
capacitor insulation was PCB (polychlorinated biphenyl), 
but because of the environmental hazards associated with 
this material, capacitors containing PCB must no longer 



be used. Owing to the dynamic operation of mining 
equipment, the correction should never be designed to 
obtain unity power factor on an average basis. If this were 
done, the resulting power factor could be leading (capaci- 
tive) for a significant portion of operation time, and the 
overall system operation could be worse than if no correc- 
tion had been attempted. The last precaution concerns 
resonance. There is a chance that undesirable resonance 
could be created if a resistance-capacitance-inductance 
(RCL) combination is formed by adding capacitance from 
the static bank to the system portion. This problem has 
been observed particularly with power centers (belt trans- 
formers) supplying solid-state belt starters. It manifests 
itself as voltage and current frequencies other than 60 Hz. 

Figure 12.29 shows the application of power-factor 
correction in a mine power center. The schematic is 
basically a reproduction of figure 12.2 with the addition of 
items 18 and 19, a static capacitor bank and its associated 
protection, respectively. Three capacitors connected in 
ungrounded wye are shown, but an integral three-phase 
capacitor with the same configuration can also be used. A 
molded-case circuit breaker affords short-circuit protection 
but also acts as a switch for removing the bank from the 
line. Sizing this breaker follows the same procedure re- 
lated earlier for other applications and can be based on 
normal capacitor current. The current can easily be calcu- 
lated from the per-phase capacity (kilovoltamperes reac- 
tive) and the system voltage. Some load centers are de- 
signed with more than one capacitor bank so that the 
magnitude of correction can be changed. 

The sizing of the capacitor is also straightforward and 
follows the basic procedures presented in chapters 3 and 4. 
For instance, consider a mining section that has been 
found to be consuming 172 kW at an average power factor 
of 0.6 lagging. The average apparent and reactive powers 
are then 



S = 



P 

pf 



172 
0.6 



= 287 kVA, 



Q L = S[sin (cos -1 pf)] = 287 (0.8) = 230 kvar. 



With a 90-kvar capacitor bank, the resultant reactive 
power and new power factor would be 

Q T = Q L - Q c = 230 - 90 = 140 kvar inductive, 



Q T 140 

Ian u-jy — -p = v7o> ^t = *-*" » 



pf = cos"^ = 0.78. 

Thus, the average power factor has been improved from 0.6 
to 0.78. As a general recommendation, power-factor cor- 
rection in mine power centers should not try to correct an 
average power factor above 0.85. 

Figure 12.29 also serves as a summary of the material 
presented on ac mine power centers. The internal compo- 
nents of an ac mine power center can be grouped into those 
associated with the high-voltage side, the transformer, and 
the utilization side. 



320 




KEY 
High- voltage coupler 
Interlock switches 
Emergency-stop switch 
Disconnect switch 
Pilot -break monitor 
High-voltage fuses 
Surge arresters 
Surge capacitors 
Power transformer 
Temperature device 
Grounding resistor 
Busway 

Outgoing circuit breaker 
Main circuit breaker 
Voltage metering 
Current metering 
Outgoing cable coupler 
Power-factor correction 
capacitors 

Circuit breaker 



To utility 

transformer 

and control 

circuit 



Figure 12.29.— Application of power-factor correction in mine power center. 



DIRECT CURRENT UTILIZATION 

Although ac utilization now dominates the mining 
industry, some segments of the mining process still remain 
better suited to the use of dc motors. The classic example 
is the use of dc series-wound motors for traction. The 
speed-torque characteristics of these motors are particu- 
larly well suited to this application, especially for locomo- 
tives and shuttle cars. This is the reason why dc utilization 
still plays a significant role in the mining process. 

There are various ways of obtaining a dc voltage for 
powering the face equipment. If a mine uses rail haulage 
at a working section, the dc equipment can be powered 
directly from the trolley feeder line. However, if a working 
section does not have direct access to a trolley feeder line, 
a rectifier must be used to convert the three-phase ac 
voltage to dc voltage. The rectifier can be a separate piece 
of equipment housed in its own enclosure and powered by 
means of a feeder cable from the ac power center, or it can 
be incorporated into a single enclosure with the ac power 
center, with the total unit being referred to as an ac-dc 
combination power center. Some mining machinery man- 
ufacturers offer face equipment with on-board rectifiers so 
the benefits of dc motors can be obtained without the need 
for a section rectifier (see chapter 14). 

Even if the rectifier is supplying a trolley, the basic 
internal components of dc power equipment remain essen- 
tially the same, and so only the dc section of a combination 
power center will be discussed. A general arrangement of 
the dc components is illustrated in figure 12.30. This 
diagram will be used as a reference to indicate the 
placement of individual components with respect to the 



y £ To control circuit 




Figure 12.30.— General arrangement of dc components for 
combination power center. 



321 



overall system. The dc circuits that are frequently associ- 
ated with trolley feeder lines will also be mentioned. 
Components and circuits that have been discussed in the 
preceding sections will be presented only when they are 
required for clarity. 

RECTIFIER TRANSFORMER 

A three-winding transformer (fig. 12.30, No. 1) is 
commonly used in a combination power center and con- 
sists of primary, secondary, and tertiary windings. Nor- 
mally, the primary winding is delta connected with the 
secondary and tertiary windings wye connected, but the 
transformer may be designed such that the primary wind- 
ing can be connected in either a delta or a wye configura- 
tion so it can be applied at two different distribution 
voltages: for example, 4,160-V delta, 7,200-V wye. For the 
latter situation, either the secondary or the tertiary wind- 
ing must be delta connected. Thus, if a resistance- 
grounded system is to be used on the delta-connected 
winding, a zig-zag or grounding transformer must derive a 
neutral. 

The full-wave bridge (fig. 12.31) is the most popular 
rectifier configuration used in combination power centers 
or section rectifiers. Here the relationship of the ac rms 
input voltage to the dc output voltage is 

V ac = 0.74 V dc . 

The nominal 300-V system voltage is the one most com- 
monly used in mining applications. Although some 600-V 
systems are still in operation, they are usually used only 
for trolley supplies. For a nominal system voltage of 300 V, 
the line-to-line ac voltage feeding the input to the rectifier 
would be 



»V in 



V f .360-Hz ripple 



Figure 12.31.— Full-wave bridge rectifier. 




Air-core 
reactors 



Main 
circuit 
breaker 



¥ ¥ ¥ 

To rectifier 

Figure 12.32.— Series reactance to reduce available short- 
circuit current. 



V ar = 0.74 (300) = 222 V. 



Hence, the voltage rating of the transformer winding 
feeding the rectifier would be 222 V. Some of the common 
ratings used in combination power center are listed in 
table 12.2. 

Table 12.2.— Typical ratings for combination power centers. 

dc, kW ac, kVA ac voltage, V dc voltage, V Input voltage, V 

75 300 480Y/277 300 7,200 

150 375 480Y/277 300 4,160 

400 480Y/277 300 7,200 

200 600 600/220 300 7,620/13,200 

300 500 480Y/277 300 7,200 

750 600/220 300 7,620/13,200 
1,000 600/220 300 7,200/4,160 

A dc is much more difficult to interrupt than an ac, 
and the amount of available short-circuit current must be 
limited to less than that of ac equipment with similar 
capacity. As a result, the impedance of transformer wind- 
ings feeding rectifier circuits is normally in the range of 
0.06 to 0.08 pu. In order to obtain this high impedance, the 
leakage reactance of the associated winding is often delib- 
erately increased during transformer fabrication. If this 
method does not provide the adequate impedance, a small 
additional impedance can be obtained by using air-core 
reactors at the output of the transformer, as shown in 
figure 12.32. However, in some circumstances, it is neces- 
sary to use a separate transformer (fig. 12.33) to limit the 



Main circuit 
breaker (ac) 




To ac 
circuits 



Figure 12.33.— Separate transformer to increase impedance 
of dc circuit. 



322 



maximum fault current to a level that can be safely 
interrupted and a magnitude that will not damage the 
i-ectifier bridge. 

A main circuit breaker (fig. 12.30, No. 2) can be 
provided to protect the transformer winding that supplies 
the dc circuit. The thermal-trip setting of this breaker is 
usually based upon 125% of the rated current of the 
winding. As an example, a winding rated at 150 kVA 
would have a full-load current of 



150 
FL " V3 (0.222) 



= 390 A. 



The thermal-trip setting would be the next rated value 
above 1.25 (390), or 488 A. If the transformer impedance is 
0.06 pu, the breaker should be capable of interrupting a 
fault current of 



If = 



390 
0.06 



= 6,500 A. 



RECTIFIER 

As seen in figure 12.30, the output of the transformer 
tertiary winding feeds the rectifier (No. 3) via the main 
circuit breaker. Again, the rectifier configuration is almost 
always a three-phase full-wave bridge. The power or kilo- 
watt requirement of rectifiers used for mining is normally 
too high to permit the use of a single diode in each leg of 
the bridge. The current rating of the rectifier bridge, and 
thus the diodes, is selected on the expected bolted-fault 
current (faulted dc output) with the power-center input 
connected to an infinite bus. This current should also 
incorporate a multiplier to account for a worst case offset. 
The rectifier should be capable of handling this current for 
the time required for the circuit breakers to interrupt 
current flow. To achieve the adequate current-carrying 
capacity, diodes are paralleled in each rectifier leg to share 
the current. 

Figure 12.34 shows a full-wave bridge rectifier with 
two diodes paralleled per leg (No. 1). The number of 
parallel diodes is determined by the individual diode 
current rating (all must be rated equally) along with the 
maximum available fault current. Each diode must also be 
rated in terms of a peak inverse voltage (PIV), which is the 
maximum voltage that can be applied across the diode in 
a reverse-biased mode without causing breakdown. For 
mine rectifier applications, the diode PIV should not be 
less than 2.5 times the nominal dc system voltage if proper 
transient suppression is used. Thus for a 300-V system, a 
rating of 800 V is commonly employed. 

Matched diode characteristics should allow sufficient 
means for sharing current when used in a parallel config- 
uration. Replacement through failure of any diode, how- 
ever, would easily upset the balance. Therefore, although 
rectifiers are supplied with matched diodes, most manu- 
facturers use current-balancing reactors (fig. 12.34, No. 2) 
to force uniform conduction of all diodes in parallel. (See 
chapter 4 for an explanation of operation.) 

Each diode is accompanied by a circuit-protection fuse 
(fig. 12.34, No. 3). The purpose of the fuse is not to protect 
the diode but to prevent a catastrophic failure to a bridge 
leg (these diodes fail in a shorted mode). The fuse current 
rating must be such that it will not initiate interruption 
for a bolted fault at the dc bus until the clearing time for 



the downstream circuit-interrupting devices has elapsed 
(that is, the time-current characteristics of the two must be 
coordinated). Each fuse has a matching light (fig. 12.34, 
No. 4) to indicate when the fuse has blown and the diode 
needs to be replaced. 

Suppression devices should always be used to protect 
the rectifier from transient overvoltages occurring from 
either the ac or dc side. Owing to their successful history 
in mining applications, selenium voltage suppressors (fig. 
12.34, No. 5) are presently the most popular devices used 
in rectifiers. They are available with rms ratings ranging 
from 30 V to 480 V in 30-V increments. The clamping 
voltage at rated discharge is approximately 2.5 times the 
rms voltage rating. For instance, the input ac voltage to a 
300-V rectifier would be 222-V rms; therefore, a selenium 
suppressor rating of 240- V rms should be selected. The 
resulting clamping voltage would be 600 V, which is 
significantly below the 800-V PVI rating determined 
earlier. Although not shown in figure 12.34, the suppres- 
sor often has each connection (five in all) protected by a 
fuse, again with a paralleled light to indicate failure. For 
rectifiers serving face equipment (150 or 300 kW), one 
selenium suppressor is used, but two or more may be 
paralleled for trolley rectifiers (450 kW or more). 

An RC-snubber circuit, as shown in figure 12.35, is 
sometimes provided to reduce commutation transients. 




Figure 12.34.— Typical full-wave bridge rectifier with two 
diodes in parallel per leg. 



Diode 



Circuit- 
protecting 
I 2 t fuse 



Hvwv- 



-a- 



dv/dt snubber 
circuit 



Indicator 
light 



Figure 12.35— Diode with RC snubber protection. 



323 



The circuit serves to reduce the voltage rate-of-rise (dv/dt) 
that might be developed across the diode. Although snub- 
ber circuits are common in mine rectifiers, some manufac- 
turers have achieved reliable diode operation without 
their use. The application of this protection technique is 
discussed in chapter 14 as dv/dt protection is required on 
thyristors. 



DIRECT CURRENT GROUND-FAULT 
PROTECTION SCHEMES 

Five basic systems of dc ground-fault protection are 
presently being used in the United States: 

1. Diode grounding, 

2. Basic grounding conductors, 

3. Relayed grounding conductors, 

4. Neutral shift, and 

5. Differential current. 

These are illustrated respectively in figures 12.36 through 
12.40. The technique used dictates the complexity of the 
circuitry that will be associated with each outlet from the 
rectifier section. Because a complete discussion is pre- 
sented in chapter 8, the systems will only be summaried 
here. The first three systems are more commonly used 
when a trolley feeder is the dc source. The neutral-shift 
and differential-current systems can only be used in con- 
junction with rectifiers serving face equipment. It should 
be noted that rectifiers for trolley systems must have both 
positive and negative conductors isolated above frame 
ground; thus, the subject of grounding is not applicable in 
that instance. 

In the diode-grounded system (fig. 12.36) the machine 
frame is tied to the grounded conductor through a ground- 
ing diode. The basic grounding-conductor system (fig. 
12.37) uses a separate conductor to provide grounding to 
the machine frame. In essence, adequate grounding-fault 
protection is not available from either system; for example, 



Machine frame 




Figure 12.36.— Diode-grounded system. 



(+) 
(-) 



-o o — ^ 



(+) 



(-) 



GND 



Machine 
frame 



ground faults in the trailing cable rely on an interrupting 
device for sensing and clearing. No additional circuitry, 
other than the grounded power conductor, is used in the 
power center. 

The relayed grounding-conductor system (fig. 12.38) 
can be sensitive to ground faults. A ground-fault relay coil 
is placed in series with the grounding conductor, and when 
a current threshold is reached, contact pickup can be used 
to trip a breaker UVR. However, parallel paths existing 
through the earth can cause fault current to partially 
bypass the relay coil, making for unreliable protection. 

The neutral-shift system is shown in figure 12.39, 
where the resistors R x and R 2 create a dc neutral point. 
Ground faults on either the positive or negative conductor 
will cause the neutral to shift. The voltage-sensing relays, 
M 1 and M 2 , detect the change. The result is sensitive 
ground-fault relaying, but the system is not selective with 
more than one outgoing circuit. 

For one type of differential-current system (fig. 12.40), 
a grounding resistor is placed between the transformer 



-o o — /- 



(+) 



=-o 



(-) 



iA 



GND 



s?=-o 



GTR 



Machine 
frame 



n 

To 

circuit breaker 

trip circuit 

Figure 12.38.— Relayed grounding-conductor system. 



To 
rectifier . 



M, ± 



To circuit breaker 
trip circuit 



(+) 



R 2 



M?) ± 



(-) 



GND 



= — o 



To circuit breaker 
trip circuit 



Figure 12.39.— Neutral-shift system. 



To circuit breaker 
trip circuit 



Grounding 
resistor 



u 




Saturable 
reactor 



\ 


1 


— < 

. i 


> — < 

i i 


1 — 1 

; 








1 

i 1 




J™ • 




1 

i i 


i 


/ 


i 



rO ► 

Ground- 
fault relay 

, (+) r 

, gndT, 



To ac 

control 

circuit 



Machine 
frame 



Figure 12.37.— Basic grounding-conductor system. 



Figure 12.40.— Differential current scheme. 



324 



neutral and load-center frame and both positive and 
negative conductors pass through a saturable reactor. The 
differential current created by a ground fault causes 
magnetic saturation of the reactor core, which in turn 
allows the ground-fault relay to pick up. The system is 
sensitive and selective and by using the technique for each 
outgoing dc circuit, practical ground-fault protection is 
available. As a result, this system is considered in more 
detail in the following discussion. 



DIRECT CURRENT CONTROL CIRCUITRY 

A control circuit using a differential-current tech- 
nique for ground-fault relaying is illustrated in figure 
12.41. A main control transformer supplies 120 V to the 
entire circuit. Heat-sensing devices are mounted to the 
heat sink for each leg of the bridge rectifier. The overtem- 
perature relay (OTR) coil is connected in series with the 
heat-sensing devices and is energized by a filtered single- 
phase bridge rectifier. The relay contacts are in series with 
the UVR of the main breaker so when a rectifier overtem- 
perature condition occurs, a sensing device opens, the 
relay resets, and the UVR trips the breaker. 

This version of differential-current relaying is sup- 
plied from the control power through a stepdown trans- 
former with a 6-V secondary. With a ground-fault current 
from 4 to 6 A, voltage is impressed cross the ground-trip 
relay (GTR) in sufficient magnitude to cause contact 
pickup. This deenergizes the UVR on the machine circuit 
breaker, which in turn interrupts the offending circuit. 



DIRECT CURRENT INTERRUPTING DEVICES 

The problem mentioned earlier concerning interrup- 
tion of dc requires the interrupting device to force the 
current to zero with an arc voltage greater than the line 
voltage. The device must be capable of withstanding the 
energy dissipated during the time that the arc exists 
across its contacts. This energy is a function of the circuit 
inductance as well as the current magnitude (3, 14). There 
are two basic interrupters used to protect dc machine 
circuits: the dc contactor and the molded-case circuit 
breaker. For trolley systems, power circuit breakers are 
employed. 

A cross-sectional view of a dc contactor is shown in 
figure 12.42. This device consists primarily of stationary 
and movable contacts, which make and break upon ener- 
gizing and deenergizing the operating coil. When the 
operating coil is energized, the armature assembly is 
attracted to the pole piece, which completes the magnetic 
circuit and causes the contacts to close. The contactor is 
designed so that the contacts close with a self-cleaning 
wiping action to ensure a good electrical contact (6). 

Arc termination is similar to that of an air-magnetic 
circuit breaker (see chapter 9). A blowout coil is located 
between the front power terminal, Z, and the stationary 
contact. The purpose of the coil is to aid in extinguishing 
the arc and to minimize burning of the contact tips. To 
accomplish this, the coil produces a magnetic field perpen- 
dicular to the arc, which causes the arc to be lengthened 
and ruptured by deflection. Since the blowout coil is 
connected in series with the circuit to be interrupted, the 
strength of the magnetic field is proportional to the 
current. A flexible braided shunt completes the connection 



■— o- 
'^mcid- 



Control 
transformer 



Lo^«- 





Hi 



OTR (49 R) 



ik il 



Rectifier 



(+) 
( 






Ti * 

-• — •- 



l| ' V . , II 



Heat-sensing devices 
dc bus 




Figure 12.41.— Representative control circuit for rectifier. 



Top 
insulator 

Lo "f / Operating x 

insulator £ oi | a Top 

plate 



Blowout 
coil 



Arc horn 



Arc shield-* 



Stationary 
contact tip 

Stationary 
contact 
assembly 



Shunt 

Spring 
washer 




Armature 
hinge pin 

Lock screw 



Movable 
contact tip 



Armature 
assembly 



Moving contact 
assembly 



Figure 12.42.— Cross section of dc contactor. (Courtesy Joy 
Manufacturing Co.) 



325 



from the movable contact to the rear power terminal, Y. 
The arc horn protects the insulation of the blowout coil 
from being burnt by the arc, and an arc shield encloses the 
contacts to confine the arc and protect the adjacent parts 
(6). 

The contactors most commonly used as interrupting 
devices for dc machine circuits have a continuous-current 
rating of 250 or 500 A. Their interrupting capability is 
approximately 10 times their continuous-current rating, 
which is significantly less than that given for molded-case 
circuit breakers. One advantage of the dc contactor is its 
simple design and ease of maintenance. In some applica- 
tions, molded-case circuit breakers are used as backup 
protection for dc contactors. It should be noted that 
molded-case circuit breakers are not rated for 600-V appli- 
cations, but dc contactors can be applied at this voltage. 

Molded-case circuit breakers that are rated at 600 Vac 
are also normally rated at 300 Vdc. The dc interrupting- 
duty rating for mine-duty breakers is either 10,000 or 
20,000 A, depending on the product. However, as with the 
dc contactor, the actual interrupting capability varies with 
the inductance of the circuit being interrupted. Labora- 
tory tests have indicated that for dc operation, the instan- 
taneous dc tripping current is about 1.3 times the 
magnetic-element trip setting for ac (14). 

Even though molded-case breakers have higher inter- 
rupting capacity, many engineers in mining operations 
have expressed concern over using them for dc applica- 
tions. The engineers relate that in practice molded-case 
devices cannot adequately interrupt dc faults, whereas 
low-voltage power circuit breakers appear to work satis- 
factorily. On the other hand, manufacturers report that 
they have had few problems with field applications of 
molded-case breakers on dc. 

In summary, the principal understanding that should 
be acquired from this power center presentation is that the 
equipment is assembled to match the mining operation 
and application. This is true whether the unit is a belt 
transformer, a trolley rectifier, or a combination ac-dc 
power center. Consequently, there cannot be a standard 
mine power center, but nevertheless, there are many 



practices recommended for mine power center construc- 
tion that must be considered in order to achieve an 
efficient operating system. 

REFERENCES 

1. Dornetto, L. D. The Importance of Grounding Systems in the 
Protection of Personnel and Equipment. Paper in Mine Power 
Distribution. Proceedings: Bureau of Mines Technology Transfer 
Seminar, Pittsburgh, Pa., March 19, 1975. BuMines IC 8694, 1975. 

2. Dutton, J. C, and W. J. McNutt. Transformer Stan- 
dards-Status and Trends. IEEE Trans. Ind. Appl., v. 16, 
Jan./Feb. 1980. 

3. Helfrich, W. J., P. M. Hall, and R. L. Reynolds. Time Con- 
stants-Direct Current Mine Power Systems. Paper in Proceedings 
of 5th WVU Conference on Coal Mine Electrotechnology (Morgan- 
town, WV, July 1980). WV Univ., 1980. 

4. Institute of Electrical and Electronics Engineers (New York). 
General Requirements for Dry-Type Distribution and Power 
Transformers. ANSI/IEEE Stand. C57.12.01-1979. 

5. Recommended Practice for Electric Power Distribu- 
tion for Industrial Plants. Stand. 141-1986. 

6. Joy Manufacturing Co. (Franklin, PA). Direct Current Min- 
ing Machinery. 5th ed., 1971. 

7. King, R. L. Development of an Electrical Engineering Course 
for Mining Engineers. M.S. Thesis, Univ. Pittsburgh, Pittsburgh, 
PA, 1977. 

8. Kline, A. D. Design Consideration of Mining Transformers. 
Paper in Conference Record -IAS 15th Annual Meeting (Cincin- 
nati, OH, Sept.-Oct. 1980). IEEE, 1980. 

9. Line Power Manufacturing Corp. (Bristol, VA). Electrical 
Power for the Mining Industry. Sales literature, undated. 

10. McGraw-Edison Co. (Canonsburg, PA). Application catalog 
240-60, undated. 

11. McGraw-Edison Co., Power System Div. (Canonsburg, PA). 
Distribution-System Protection Manual. Bull. 71022, undated. 

12. Ohio Brass Co., Ensign Electric Div. (Mansfield, OH). Mine 
Power Equipment. Sales literature, BHE37, undated. 

13. RTE Corp. (Waukesha, WI). Transformers. Undated. 

14. Shimp, A. B., and D. A. Paice. Application of Molded-Case 
Breakers on DC Electrical Systems in Coal Mines. Paper in Mine 
Power Distribution. Proceedings: Bureau of Mines Technology 
Transfer Seminar, Pittsburgh, Pa., March 19, 1975. BuMines IC 
8694, 1975. 



326 



CHAPTER 13.— SWITCHHOUSES AND SUBSTATIONS 1 



Switchhouses, substations, and power centers comprise 
the major power equipment in mine power systems. Chapter 
12 covered the design and construction of mine power cen- 
ters, and this chapter will follow a similar format to present 
switchhouse and substation arrangements. 



SWITCHHOUSES 



associated control circuitry for figure 13.1 is given in figure 
13.2. The operation of this circuit will be discussed later. 

As in mine power centers, interlock switches are 
positioned around the side covers and top covers of the 
switchhouse and are wired into the incoming pilot circuit 
to trip the upstream breaker in the event that a cover is 
removed. An emergency stop button is also provided. 



Switchhouses are contained in metal-clad enclosures, 
similar in construction to those of power centers. This power 
equipment consists of visible disconnects and sectionalizing 
units, although the term switchhouse is normally equated 
with just the sectionalizing equipment. The principal func- 
tion of the disconnect switch is to remove power manually 
from downstream mine power equipment that is connected 
to distribution. As in power centers, it must be possible to 
determine the position of the disconnect (load break) switch 
visually through a window in the enclosure wall. In under- 
ground coal mines, the power removal function must be 
located within 500 ft of the point where power enters the 
mine, and additional disconnects are incorporated into all 
power equipment that receives power from high-voltage 
distribution. This disconnect-switch equipment can be de- 
scribed simultaneously with switchhouses. The high-voltage 
side of the power center that was shown in figure 12.2 
contained the same components. 

The prime role of the switchhouse is to provide pro- 
tective relaying in the distribution system and to allow 
branching of the radial system. The principal component 
is an automatic circuit breaker. The equipment name is 
modified depending upon the number of circuit breakers 
and associated protected outgoing circuits: for example, a 
double switchhouse contains two breakers and circuits. 
Because of size limitations, the units in underground mine 
power systems are rarely larger than double switchhouses, 
but surface mines often incorporate four-breaker switch- 
houses, or switching skids as they are commonly called. 
The schematic diagram for surface or underground appli- 
cations is practically the same: the only basic difference is 
the repetition of internal components to correspond to the 
number of circuit breakers used. 



SWITCHHOUSE INTERNAL COMPONENTS 

A general arrangement for a single-breaker switch- 
house is illustrated in figure 13.1. Incoming high-voltage 
power enters the switchhouse through the input receptacle 
to the load-break switch and the feedthrough receptacle. 
Some circuit-interrupting devices have disconnect switches 
incorporated in their construction, as shown in this diagram; 
others have a load-break switch located on the line side of the 
breaker. In either case, the switch must have an external 
operating handle with a mechanism for locking the switch in 
the open position. The output (load side) of the interrupting 
device feeds the branch receptacle. The current-sensing de- 
vices for the protective-relaying circuits are usually situated 
between the breaker and the branch receptacle, while the 
control transformer is located on the incoming side. The 



1 The author wishes to thank Thomas Novak who prepared original 
material for many sections of this chapter. 





Branch x 
receptacle _± 

To line ^ ground- 
overcurrent check 

relays and monitor 

ammeter 



Cover Emergency 
interlocks stop 



Figure 13.1.— Diagram for typical single switchhouse. 




I — ww- | 

■ — WW- 

" — WW-' 
o 




-tpo ^/rm,^ 



52 a 



52 b 




i TgcsT 



GCS 



»To pilot 

— o — 

UV52 



"UUUU 



Heater 
strips 







T 52a 

Q 52 trip coil 



Battery 



KEY 
I Instantaneous contact 
T Timed contact 
SI Seal- in contact 



K§h 



ISsT 



!-<§>. 



'si 



X^ 



-W- 



X-%- 



GCS _i_ Trip button 
o o — 



50/5H 
50/51-2 
50/51-3 

51-G 



Figure 13.2.— Control circuitry for single switchhouse using 
battery tripping. 



327 



A typical arrangement of a double switchhouse is 
presented in figure 13.3, and its control circuitry is shown 
in figure 13.4. The figures again show breakers with 
incorporated switches, and the repetition of circuitry men- 
tioned earlier is obvious. In addition, the circuits in 
branch A work independently of branch B, and vice versa. 
Where the breakers do not have built-in switches, one 
load-break switch is usually located upstream from both 
breakers, and serves as a disconnect for all branch circuits, 
but not the feedthrough. 

The incoming circuitry to the circuit breakers, shown 
in figures 13.1 and 13.2, is nearly identical to the incom- 
ing circuitry to the power center transformer discussed in 
chapter 12 and shown in figure 12.2. Surge arresters and 
switch prebreak monitors are not incorporated in the 
switchhhouse schematics, but they are often found in 
practice and are recommended. A typical schematic for the 
disconnect switch (the equipment, not the component) 
would be similar to figure 13.1 if the circuit breaker and 



Input 
receptacle 




Cover Emergency 
interlocks stop 



^rrrru 



Branch A 
receptacle 



rtt 



To ground- To phase 
trip relay overcurrent 
A relays-A 




JT TTl 



Branch B 
receptacle 



-, -rrm - 



To ground- To phase 
trip_relay overcurrent 
relays-B 




To control 
circuits AandB 



DC, 

Figure 13.3.— Diagram for typical double switchhouse. 



Control 
transformer 



I — VWV— | 

(i — v\M—-<> str 'P 
, , .... . , heaters 



52 A/a 

Hi — 



52A/b 



Closed 
'M/W\rt 




-© ■• 



50/51-3A 



51-GA 



■SI 



(SI) 



fa- 



— i¥- 

GCS-A 




KEY 
I Instantaneous contact 
T Timed contact 
SI Seal -in contact 



Trip button -A Trip button -B 

Figure 13.4.— Control circuitry for double switchhouse using capacitor tripping. 



328 



current transformers (CT's) were removed. This similarity 
is important in mine power-equipment design because 
when circuits serving identical functions have the same 
arrangement and construction, the maintenance crews 
can work on all equipment containing these circuits with- 
out surprise. This both simplifies mine training and 
increases the safety factor for maintenance personnel. 



SWITCHHOUSE PROTECTIVE RELAYING 

Induction-disk overcurrent relays are used for most 
protective-relaying applications in switchhouses. General 
relay operation and characteristics have already been 
discussed in chapter 9. For mining applications, inverse- 
time characteristics are usually employed for line- 
overcurrent relaying, while very-inverse-time relays are 
commonly used for ground-fault protection. 

These relays have a range of adjustments so they can 
be applied to a variety of situations: the operating time 
can be controlled by the time dial setting; pickup current 
is adjusted by the main coil taps. Since the resulting time- 
current characteristics are the same for each tap, curves 
are plotted in terms of multiples of the pickup value. A 
typical family of curves for an inverse-time relay is given 
in figure 13.5 (ll). 2 An instantaneous relay can be incor- 
porated with the induction-disk (timed) relay, which will 
respond to the same actuating quantity. The actuating 
currents that are delivered to overcurrent relays are ob- 
tained by CT's. The CT's provide insulation from the 
high-voltage circuit and supply the relays with currents 
that are proportional to the currents in the line conduc- 
tors, but sufficiently reduced. When applying a CT, critical 
consideration should be given to transformer burden and 
performance. Chapter 10 can be consulted for information 
on necessary calculations. 

The pickup of an overcurrent relay should be selected 
so it will operate for all short circuits within its primary 
zone of protection and provide backup protection in adjoin- 
ing zones. Figure 13.6 will be referred to as an example. 
Breaker 1 should provide primary protection for line 1. At 
the same time, its relays must be coordinated with the 
relays of breakers 2 and 3 to afford backup protection for 
lines 2 and 3. To ensure selectivity under all circum- 
stances, the pickup value of any given relay should be 
somewhat higher than the pickup of downstream relays. 
The time delay of overcurrent relays must also be adjusted 
to provide selectivity with the relays of the immediately 
adjoining zones. 

To provide proper selectivity, a bolted three-phase 
fault can be assumed as a basis for adjusting line- 
overcurrent relays, and a bolted line-to-ground fault would 
form the basis for the ground-relay setting. Selectivity 
under these conditions should ensure selectivity at lower 
currents. 

For the fault shown in figure 13.6, the relay at breaker 
2 must close its contacts, and breaker 2 must trip and 
interrupt the flow of short-circuit current before the relay 
at breaker 1 closes its contacts. The calculation must also 
account for the overtravel of the relay at breaker 1. The 
time delay of breaker 1 that is needed to provide selectiv- 
ity with breaker 2 can be determined by 



i \ \ i i i | — i — i — i i i i i i i i i i i 



Ti = T 2 + B 2 + C-! + F, 



(13.1) 




1 ''^L J. 1. '_ ' ' ' UU I I I 






15 2 3 4 5 6 7 8 9 10 15 20 

MULTIPLE OF PICKUP 
Figure 13.5.— Typical family of curves for inverse-time relay. 



Line I . — , . . Line 2 

HZr¥ ^ 



1 



Fault 



Line 3 



2 Italicized numbers in parentheses refer to items in the list of references 
at the end of this chapter. 



Figure 13.6.— Illustration of fault location for adjusting 
selectivity. 



where T 1 = operating time of relay at breaker 1, s, 
T 2 = operating time of relay at breaker 2, s, 
B 2 = short-circuit interrupting time of breaker 2, 

s, 
O x = overtravel time of relay at breaker 1, s, 

and F = factor of safety, s. 

As was shown in chapter 10, the sum of B 2 , 1? and F can 
normally be assumed to be 0.4 s. When coordinating the 
time delays of inverse-time relays, the process should start 
with the most downstream relays and work back toward 
the substation. 

The typical connections for the line-overcurrent relays 
and CT's used in switchhouses are illustrated in figures 
13.2 and 13.4. The inverse-time units are labeled 51, while 
the instantaneous units are labeled 50. Figure 13.2 shows 



329 



three wye-connected CT's driving three wye-connected 
overcurrent relays, whereas figure 13.4 shows two open- 
delta-connected CT's driving two open-delta-connected 
overcurrent relays. Although protection for all three lines 
can be obtained with the open-delta connection, this 
approach is neither as accurate nor reliable as the wye 
connection. 

The best ground-fault protection is obtained through 
zero-sequence relaying, which is shown in both figure 13.2 
and figure 13.4. The CT ratio is typically 25:5 or 50:5. 
Unlike the power-center applications in chapter 12, the 
secondary current from the CT, which has an induction- 
disk relay as burden, is proportional to the ground-fault 
current as dictated by the turns ratio of the CT. However, 
because of the necessarily large CT windows, burden 
matching and actual testing are essential to obtain reli- 
able sensitive relay pickup. Here the aim should be to 
obtain 50% of the current limit. 

Residual ground-fault relaying can be found in some 
switchhouses, but adequate sensitivity is usually not 
available with this system because of errors caused by CT 
saturation and unmatched characteristics. Some success 
in eliminating this problem has been achieved with the 
use of solid-state relays, as discussed in chapter 14. 



must have a close-and-latch rating, based on asymmetrical 
current, that is greater than 1.6 times the maximum 
symmetrical fault current. 

Sometimes an automatic load-break switch inter- 
locked with fuses (as discussed in chapter 12) replaces the 
power circuit breaker in the most downstream switch- 
house in the distribution system. A small delay is built 
into the switch tripping circuit to remove the problem of 
the switch's trying to clear fault currents. The time delay 
is coordinated with the maximum period that a second 
fuse takes to clear a line-to-line or three-phase fault. 
Manual interlocking with the fuse elements is therefore 
not applicable. The only automatic load-break switch 
suitable for this approach is the type where the fuses have 
elements that activate a contact set upon separation and 
can be used to power the tripping mechanisms in conjunc- 
tion with the control circuitry. 

For additional information, note that chapter 9 con- 
tains extensive information about different power circuit 
breakers and their advantages and disadvantages, and 
chapter 10 covers the selection of continuous-current, 
interrupting, and close-and-latch ratings for these devices. 



SWITCHHOUSE CONTROL CIRCUITS 



POWER CIRCUIT BREAKERS 

Three types of power circuit breakers can be found in 
mine switchhouses: live-tank oil (or minimum oil), dead- 
tank oil (OCB), and vacuum (VCB). Live-tank oil circuit 
breakers have been plagued with maintenance problems, 
the most outstanding being inadequate oil levels after five 
or fewer interruptions. This requires extremely frequent 
inspections, which are impractical for most mining opera- 
tions. Three-phase dead-tank OCB's in 250- to 500-MVA 
interrupting capacities are used exclusively by many min- 
ing companies with tremendous success. Despite this, 
their use has been substantially curtailed in recent years 
because of cost, availability problems, polychlorinated 
biphenyl (PCB) contamination, and their large size for 
thin coal seams. In fact, some States prohibit the use of 
OCB's in underground applications above 10 kV. VCB's 
have been criticized in terms of the transient overvoltages 
that their high efficiency can create. However, as demon- 
strated in chapter 11, these problems can usually be 
related to poor power-system design practices. When rea- 
sonable precautions are met, the VCB is perhaps the most 
effective high-voltage power circuit breaker compared 
with other types that have equal voltage and interrupting- 
capacity ratings. It is by far the most popular circuit 
breaker used in switchhouses. 

The voltage ratings of power circuit breakers corre- 
spond to the insulation class used in the mine distribution 
system, but manufacturers offer VCB's rated at 15 kV, 
which are applicable for 4.16-kV through 14.4-kV nominal 
voltage systems. Continuous-current ratings of 400 and 
600 A are the most common in switchhouse applications. 
Since high-voltage circuit breakers begin interruption a 
few cycles after the first-cycle fault current peak, the 
interrupting rating of these breakers is based on symmet- 
rical fault current. (Many power breakers are also rated on 
an asymmetrical interrupting-current basis.) A 400-A 
VCB has a typical interrupting rating of 4, 000- A symmet- 
rical, with the 600-A model rated at 12,000 A. The breaker 
must also be able to withstand the physical stresses 
resulting from first-cycle fault currents. Thus, the breaker 



Most control-circuit schemes found in switchhouses 
have a similar operation. Thus, for convenience of discus- 
sion, the typical control shown in figure 13.4 is reproduced 
in figure 13.7 and the control shown previously as figure 
13.2 is repeated as figure 13.8. Figure 13.7 illustrates the 
control circuit of a double-breaker switchhouse. Each 
circuit breaker has its own independent control circuit. 
Control circuit B is an exact duplicate of control circuit A; 
therefore, only circuit A will be discussed. 

The actuating currents for the line overcurrent relays 
(50/51-1A and 50/51-3A) are obtained by the open-delta- 
connected current transformers. The relays are also con- 
nected in an open-delta configuration through the amme- 
ter and ammeter switch. The ground-fault relay 51-GA is 
used in a zero-sequence ground-trip circuit. 

A single-phase control transformer steps down the 
distribution voltage to 120 V to supply both control 
circuits and the strip heaters. The strip heaters are 
resistive devices placed in the switchhouse enclosure to 
provide heat that minimizes moisture accumulation. The 
red lamp is connected in series with the normally open 
(NO) auxiliary contacts (52 A/a) of the circuit breaker, 
while the green lamp is in series with the normally closed 
(NO auxiliary contacts (52 A/b). The auxiliary contacts 
are mechanically controlled by the opening and closing 
mechanism of the breaker. When the breaker is closed, 52 
A/a also closes and causes the red lamp to give a visual 
indication that the breaker is closed. At the same time, 52 
A/b opens, which causes the green (open) lamp to be 
extinguished. The reverse procedure follows when the 
circuit breaker is opened. 

The ground-check system of branch A (GCS-A) also 
obtains its power from the control circuit as shown. The 
output of the GCS is connected to the pilot and grounding 
conductors of the outgoing cable of branch A to ensure 
continuity of the downstream grounding conductor. 

A capacitor-trip device is used to supply power to the 
circuit breaker trip circuit in the event of a fault or 
overcurrent condition. A diode is used as a half-wave 
rectifier to charge the capacitor (C). The capacitor stores 
energy to provide reliable tripping of the breaker and 



330 



minimize nuisance trips. The voltage-sensing relay (CR) 
operates when the capacitor is fully charged and causes 
the lamp to provide a visual indication of a full charge. 

Another NO set of auxiliary contacts (52 A/a) is placed 
in series with the trip coil of the circuit breaker to 
disconnect the tripping circuit from the control voltage 
when the breaker is open. The contacts of the line- 
overcurrent relays and the ground-fault relays are con- 
nected in parallel as shown. If either the instantaneous or 
the inverse-time unit of any of these relays is actuated, its 
associated contacts close, which completes the path and 
allows the capacitor to discharge through the trip coil. 
When the trip coil becomes energized, it causes the trip 
mechanism of the breaker to release and open the breaker. 

A push button and an NC set of ground-check contacts 
(GCS-A) are also parallel with the contacts of the overcur- 
rent relays. The GCS-A contacts should remain open, 



provided that the control circuit is energized and the 
integrity of the associated pilot and grounding conductor 
is maintained. If the continuity of the loop circuit is lost, 
the GCS-A contacts will close, which results in tripping 
the breaker. Depressing the push button provides a conve- 
nient means of tripping the breaker, as well as a means of 
testing the operation of the capacitor trip device. 

The control circuit of figure 13.8 is similar to the 
above control circuit, with some minor exceptions. Three 
CT's and three line-overcurrent relays are connected in a 
wye configuration to provide a more precise and reliable 
operation. A battery, which is charged by a full-wave 
bridge rectifier, is used as the energy storage device, 
rather than a capacitor. An undervoltage release (UVR) 
(UV 52) is also included, with an NO set of ground-check 
contacts in series with it. 



Control 
transformer 




nnm 



I — WW— i 
<| — w*—>> 

ii — WW—" 



Strip 
heaters 




50/51-1A 



50/51-3A 



51-GA 



i SI 



(siV* 



,SI 



CSD-* 



.SI 



GCS-A 



fa 



52B/a 



52B/b 



Closed 




Open 



1 f 



GCS-B 



2_f 



To pilot 
B 



ii WW- 



-@- 



o 



^52B/b 

^52B 
V/ trip coil 

I 



CR 



50/51-1 B 



50/51-3B 



51/GB 



i SI 



(SI 



.SI 



.SI 



GCS-B 



R°fl 



■PPF 



AMSW 



KEY 
I Instantaneous contact 
T Timed contact 
SI Seal -in contact 



Tripbutton-A Trip button-B 

Figure 13.7.— Typical control circuit for double switchhouse using capacitor tripping. 



331 




I — WW— i 

, — ww- 

' — WW-' 
o 




^p ^jrrr^ . 



52 a 



52 b 




o r.rc o 



GCS 



GCS 



»To pilot 

— o — 

UV52 



"UUUJ 



Heater 
strips 




- 1 - 52 a Battery 

5 52 trip coil 



KEY 

I Instantaneous contact 
T Timed contact 
SI Seal-in contact 



* 



'si 



K^ 



x_^ 



l-@- 



-H^ 



i-^ 



GCS _i_ Trip button 
o o — - 



50/51-1 
50/51-2 
50/51-3 

51-G 



Figure 13.8.— Typical control circuit for single switchhouse 
using battery tripping. 



SWITCHHOUSE DESIGN 

At first glance the design of switchhouses is not 
nearly as complicated as that of mine power centers, but 
the simpler circuitry can be misleading. Incorrect sizing or 
adjustment of the internal components can have a disas- 
trous effect on the mining operation. Perhaps the most 
serious concern is coordination, because it is in switch- 
houses that most adjustments must be made to obtain 
sensitive and selective protective relaying. For this reason 
alone, load-flow and fault analysis of the entire mine 
power system must be available. Comparison of these data 
allows the judicious selection of the required transformer 
ratios, relay time-current characteristics, pickup ranges, 
and device ratings. Some additional comments about the 
process are in order. 

Because of the need for a uniform design, it is best for 
all switchhouses to have an identical assembly. In other 
words, all single switchhouses should have the same 
components, and one section of a multiple-breaker switch- 
house would be basically equivalent to a single switch- 
house. This may not be economically possible in some 
mines and not practical in others, yet there would be a 
specific advantage: switchhouses could be placed any- 
where in the mine and then adjusted to suit that location. 
This would benefit any major movement of equipment 
within the complex. 



To provide such uniformity, all power circuit breakers 
should be the same, with 

1. The continuous-current rating sized to the maxi- 
mum continuous current through any switchhouse in the 
mine (demand factors should be applied), 

2. The interrupting-current rating selected to stop the 
maximum short-circuit current at any location in the 
distribution system (opposed to the preceding statement, 
the required rating here might be greater than symmet- 
rical rms short-circuit current, depending upon the time to 
interruption), and 

3. Close-and-latch rating sized to the maximum asym- 
metrical first-cycle rms fault current (convention is to use 
1.6 times the interrupting-current rating). 

Load-break switches need this same close-and-latch rat- 
ing, rather than a multiple of their interrupting duty. The 
ampacity of all power conductors within the equipment 
would be sized and braced accordingly, but the size should 
never be less than 4/0. 

For overload and short-circuit protection, multiratio 
CT's might be necessary, but experience has shown that no 
more than one or two types are required for the majority of 
applications: for example, 150/300/600:5 A, 300/600:5 A, 
75/150/300:5 A, and 150/300:5 A are common ampere-turns 
ratios. With precise burden matching and accuracy calcula- 
tions, common induction-disk relays can work reliably and 
selectively. Obviously the pickup range for the timed and 
instantaneous elements must correspond to the entire range 
needed for coordination. One range for each element might 
be inadequate for the mine's needs, but the relays are easily 
interchangeable, and one or two spare sets of relays can be 
maintained at the mine to meet all demands of line- 
overcurrent protection. Correct burden and accuracy must be 
verified for any relay-CT combination. 

Adequate ground-fault protection, however, can be a 
difficult goal to attain. As stated earlier, burden and 
accuracy matching become critical. Yet even with precise 
matching, reliable pickup at a desirable ground-fault 
current threshold sometimes cannot be achieved. The 
large window required on the zero-sequence CT introduces 
induction problems that can cause inadequate secondary 
current at rated burden. The main result is that both the 
timed and instantaneous elements of induction-disk relays 
cannot be used simultaneously for ground-fault protection 
with a 50:5 or even a 25:5 CT ratio. 

A solution to this ground-fault problem is to use only 
timed elements and apply the minimum time dial setting 
to the most downstream switchhouse. The radial distribu- 
tion arrangement can then be established such that no 
more than five relays exist between the substation trans- 
former secondary and the extreme end of distribution (see 
chapter 10 for discussion). This enables practical ground- 
fault coordination to be achieved from one tap-setting 
range and only one relay type. However, such problems as 
tolerances in off-shelf components can still remain. The 
best recourse is to apply testing that simulates a ground 
fault through the CT, then observe the response of the 
induction-disk relay. The optimum location for this testing 
is at a switchhouse in place within the mine power system. 
During initial equipment installation, the tests are not 
difficult, but subsequent maintenance checks may be 
impractical. As an alternative, it is possible to include a 



332 



test circuit, such as shown in figure 12.20, in the switch- 
house circuitry for each ground-fault relay. The resistor 
and wire turns about the CT would then be adjusted for 
the required relay pickup; for instance, 2 A and 6 turns 
would simulate a 12-A ground-fault current or 48% pickup 
on 25-A ground current limit. 

A final point must be made about switchhouse con- 
struction. Reference has been made to surge capacitors in 
chapters 11 and 12: in the past, some manufacturers and 
operators incorporated these devices to increase the sys- 
tem's characteristic impedance in an attempt to limit 
transient overvoltages created by VCB chopping. However, 
the use of discrete capacitors of any kind on the load side 
of a vacuum interrupter, including power-factor correction 
types, can set the stage for prestrike and capacitance- 
switching transients. Hence, such devices should never be 
specified within a switchhouse. (A surge capacitor used 
directly across transformer and motor windings to limit 
the rate of voltage rise is a feasible means of protection if 
needed.) For similar reasons, switchhouse and power- 
center components must never be combined into a single 
enclosure, for instance, by using a VCB in place of the 
high-voltage fuses that were shown in figure 12.2. Con- 
versely, surge arresters are desirable transient protection, 
and distribution-class types should be installed in all 
switchhouses. 



ever possible. A substation is used to receive the utility 
power before distribution to the mine and other facilities. 
This complex of equipment contains switching and protec- 
tive relaying, establishes the grounding system, and may 
or may not involve transformation to a distribution volt- 
age. The substation design should be based on personnel 
safety, reliability of operation, and ease of maintenance, 
and limit damage from fire, lightning, and equipment 
malfunction (2). The utility supply voltage for mines may 
range from 230 to 13.2 kV and less, depending upon utility 
company availability and the economics and location of 
the mine. 

The substation components are often mounted on 
concrete pads. Power-conductor terminal structures, com- 
monly made of galvanized steel but also of wood, support 
the conductors and cables that provide connections for 
transformers and circuit breakers. Insulators, usually 
made of glazed porcelain or Pyrex heat-resistant glass, 
insulate the overhead conductors from the supports. An 
overall view of a typical mine substation is given in figure 
13.9. The main substation is usually a permanent instal- 
lation, but portable or unit substations are becoming very 
popular for small load requirements. 



BASIC SUBSTATION ARRANGEMENTS 



SUBSTATIONS 

It is common practice in the mining industry to 
purchase electrical power from a utility company when- 



The most popular distribution system used in mining is 
expanded radial A simplified sketch of this system is illus- 
trated in figure 13.10, with the substation noted. Radial 
techniques have the lowest initial cost since a single trans- 
former supplies all mine circuits and there is no duplication 








, ■ J :.,_ ...' * ■ M'4'-~ 4-" ~f\ . At 



wi V? * 



Figure 13.9.— Overall view of main substation serving mine. 



333 



Utility 



Utility 




Figure 13.10.— Radial distribution applied to underground 
mine and its surface facilities. 



of equipment. Operation and expansion is simple, but this 
simplicity leads to its major disadvantage: if a major compo- 
nent fails, the entire system can be effected. Nevertheless the 
logistics of mining, especially underground mining, often 
dictate the use of radial systems. 

Single-Ended Substations 

The substation design employed for radial systems is 
often termed single-ended. Figures 13.11 and 13.12 are 
one-line diagram examples, and the only difference be- 
tween the two is found on the primary side of the trans- 
former. The substation components, much like those con- 
tained in power centers, can be grouped into three general 
sections. The primary section provides connection with 
one or more incoming utility lines, switching and inter- 
rupting devices, and transformer protection. The trans- 
former portion may include one or more transformers, a 
neutral point to establish grounding, and automatic tap 
changing, if used. The secondary section or distribution 
side provides for the connection of one or more secondary 
feeders, each having switching and interrupting devices. 

Transformers (primary side) may be protected by 
high-voltage power fuses (fig. 13.11) or by a circuit breaker 
and associated overload and short-circuit relaying (fig. 
13.12) or, less frequently, by both. Fuses or a circuit 
breaker is usually all that is necessary. Ground-fault 
protection could also be provided on the primary (not 
shown), but this relaying is restricted to circuit breaker 
applications and its inclusion depends to some extent on 
utility company requirements. The sizing of components is 
similar to that already discussed for power centers and 
switchhouses. 

The outgoing-circuit protective relaying covers ground 
faults, overloads, and short circuits, and establishes a 
primary protection zone as far as the first downstream 
switchhouse. There are two recommended backup ground- 
fault relays: neutral-current sensing between the trans- 
former bushing and the top of the grounding resistor, and 
potential relaying about the grounding resistor. The first 
ensures backup if the resistor shorts, while potential 
relaying provides protection against open-resistor hazards. 



Station y\ 
ground ^~ 



Surge 

arrester 

— o o — 



? Gang-operated 
switch 

I Power fuse 



A uLu Power 



i— o 



51-N 



52 



transformer 

-3 pt 

Circuit 
breaker 



59-G 



50/51 




T 

Pilot 



. Disconnect 
/ switch 

Surge 

arrester 

— o o — 



^Station 
"V* ground 



Power 
conductors 



To borehole 



Grounding 
conductor 

Figure 13.1 1.— One-line diagram for single-ended substation 
with fuse-protected transformer. 



59-G 




.v. Station 
\s* ground 



vy Station 
^ ground 

* 

Power 
conductors 

Grounding 
conductor 

Figure 1 3.12.— One-line diagram for single-ended substation 
with circuit-breaker-protected transformer. 



334 



Both problems can occur in this surface installation be- 
cause of exposure. The ground-fault protection must be 
coordinated with that in the downstream switchhouses. 
Line-overcurrent relays must coordinate with downstream 
protection as well as with the transformer primary protec- 
tion and the utility. Circuit breaker, relay, and transformer 
sizing and selection generally follow the procedures given 
earlier, this time closely related to switchhouses. 

The substation has two separate ground beds: the 
station (or system) and the safety (or mine). The grounding 
resistor is tied to the safety bed, and a grounding conduc- 
tor extends into the mine. Ground-check monitoring of 
this conductor is required. The surge arresters, substation 
fencing, and other metallic parts are grounded to the 
station ground bed. 

All substation components should be located within a 
fence posted with danger signs; the internal area, includ- 
ing the fence, is called the substation area. Specific 
concerns about substation components will be given in the 
following sections. 



SUBSTATION TRANSFORMERS 

Transformers for permanent surface substations are 
almost always liquid (oil) immersed and built to IEEE 
standards (5). A typical substation transformer is shown in 
figure 13.14. Capacities commonly range from 5,000 to 
30,000 kVA (sometimes less, but rarely more). In addition 
to the mine, the loads may include a preparation plant and 
general surface loads, such as pumps, ventilation fans, 
maintenance shops, and office and bathhouse facilities. 
When an underground mine is involved, it is recom- 
mended that a separate transformer be used for the 
underground power system. 

The selection of transformer capacity must be based 
on estimation of electrical load. Up to 2,000 or 3,000 kVA, 
a general rule of 1 kVA/hp of connected load is perhaps 
satisfactory. This provides more than adequate capacity 






Utility 



Utility 



Double-Ended Substations 

To enhance the reliability of the distribution circuit, 
some mines utilize a secondary-selective system. Two 
substation secondaries, as shown in figure 13.13, are 
connected with an NO tie circuit breaker. The combination 
is commonly but not always in the same substation area. 
Both substation halves are identical, and each is basically 
a single-ended unit but often with only one station and one 
safety ground bed. Under normal operation, each substa- 
tion independently feeds 50% of the load, and the distri- 
bution system on either side is actually expanded radial. If 
a primary feeder or a transformer fails, the main second- 
ary breaker on the inoperative circuit is opened, and the 
tie breaker is closed either manually or automatically. 
With automatic operation, relaying between the second- 
ary breakers and the tie breaker must be interlocked so 
that the tie breaker will not close on a faulted distribution 
system. For instance, if the main secondary breaker is 
tripped by its associated relaying, the same relays should 
prevent tie breaker closure. 

To permit continued operation with one transformer, 
consideration must be given to the following (8): 

• Oversizing both transformers (instead of 50%, 80% 
of load) so one can temporarily carry the total load (at 
125% of its capacity), 

• Providing forced-air cooling to the transformer that 
remains in service during the emergency period, or 

• Oversizing both transformers so that one can carry 
the total load. 

These substations, also called double-ended, can be eco- 
nomical when the total substation capacity is above 5 
MVA. Two separate incoming utility lines maximize the 
advantage of secondary-selective over radial systems. 
However, because of high transformer reliability, the eco- 
nomics of using a double-ended substation with one incom- 
ing line is considered questionable by some. Yet a definite 
advantage of this technique over single-ended types is that 
maintenance chores can be performed without power out- 
age, such as insulator cleaning, component adjustment 
and replacement, and no-load tap changing. 



I 

Mine 
loads 



Tie 
circuit 
breaker 



Mine 
loads 



Figure 13.13.— Simplified one-line diagram for double-ended 
substation. 




Figure 13.14.— Typical liquid-immersed transformer in 
substation. 



MB^MM 



335 



for the load and will allow minimal load growth. Note that 
a worst case demand factor of 0.75 to 0.85 is built into the 
selection. Such oversizing will also allow for an adverse 
power factor. 

When the requirements are greater than this, a more 
precise load estimate is required if optimum equipment 
costs are to be achieved. Considering the common capacity 
needs for the majority of mining operations, a load-flow 
analysis must be performed. The estimation may be made 
by using the demand factors discussed in chapters 4 and 8. 
In some cases, capacity might be determined on a per- 
connected-kilovoltampere basis (about 0.75 per connected 
horsepower) to allow for the growth in demand factor that 
always exists. 

Table 13.1 shows recommended impedance values on a 
percentage basis for a liquid-immersed transformer (5). 
These values cover a capacity range from 500 to 30,000 
kVA and are based on primary voltage, whereas mine 
power-center values are based on capacity. To provide a 
reasonable limit to short-circuit currents, the reactance 
should never be less than 5.0% (0.05 pu). 

Past practice in the industry was to use three single- 
phase transformers for easy replacement of failed units, 
and a spare was often included to facilitate the changeover. 
However, improved manufacturing techniques have now 
resulted in transformers' being among the most reliable 
components; thus, the common practice today is to install 
three-phase integral units (13). The same advantages 
given for power centers hold for substations. 

The standard ratings for substation liquid-immersed 
transformers are based on an allowable average winding- 
temperature rise of 65°C. A rating related to a 55°C rise 
is also available but is no longer listed as a standard (5). A 
transformer rated with the 65 °C rise can be loaded to 
112% of the 55°C-rise specification. The transformer ca- 
pacity always involves a self-cooled [over-air (OA)] rating 
but may also have a forced-air-cooled rating (FA). Standard 
capacities for OA-rated and FA-rated substation transform- 
ers are listed below (8): 



Secondary 


Secondary 


< 1,000 V 


2,400 V and above 


5.75 


5.5 


6.25 


6.0 


6.75 


6.5 


NAp 


7.0 


NAp 


7.5 


NAp 


8.0 



Table 13.1.— Standard impedance for liquid-immersed 
three-phase transformers, percent 



Primary voltage, V 

2,400 to 22,000 

26,400 to 34,500 

43,800 

67,000 

115,000 

138,000 

NAp Not applicable. 



SUBSTATION SWITCHING APPARATUS 

Dead-tank OCB's, such as those illustrated in figure 
13.15, are the most common interrupters used in substa- 
tions. In recent years, air-magnetic, vacuum, and sulfur 
hexafluoride (SF 6 ) circuit breakers have gained in popularity 
because of their reduced and simple maintenance, but for 
applications at utility voltages of 69 kV and higher, OCB's or 
SF 6 breakers are employed almost exclusively (13). SF 6 
circuit breakers use a fluorocarbon gas for the arc- 
interruption medium, and their high cost usually restricts 
their use at 69 kV and above. On the other hand, the use of 
OCB's at lower substation voltages often results in lower 
costs and easier installation than the alternatives. 

Because of the transformer overload capacity and the 
lack of overload capacity in circuit breakers, main break- 
ers of the transformer secondary should have a continuous- 
current rating 25% greater than the anticipated top 
continuous-current rating of the transformer. Obviously, 
the voltage, interrupting-current, and close-and-latch rat- 
ings must be sufficient for the system to which the breaker 
is applied. These demands may be greater than for inter- 
rupters in switchhouses. 

Reclosers 



55°C OA: 






500 


2,500 


12,000 


750 


3,750 


15,000 


1,000 


5,000 


20,000 


1,500 


7,500 


25,000 


2,000 


10,000 


30,000 


65 °C OA: 






560 


2,800 


13,440 


840 


4,200 


16,800 


1,120 


5,600 


22,400 


1,680 


8,400 


28,000 


2,240 


11,200 


33,600 


55°C FA: 






575 


3,125 


16,000 


862 


4,687 


20,000 


1,150 


6,250 


26,667 


1,725 


9,375 


33,333 


2,300 


12,500 


40,000 


65°C FA: 






644 


3,500 


17,920 


966 


5,250 


22,400 


1,288 


7,000 


29,867 


1,932 


10,500 


37,333 


2,576 


14,000 


44,800 



Oil circuit reclosers have proved to be extremely reliable 
as interrupters on the transformer secondary at distribution 
voltages to 15 kV (13). A circuit recloser is a circuit breaker 



If the transformer is so equipped, the unit could also have 
a forced-oil-and-air (FOA) rating. 




Figure 13.15.— Dead-tank OCB in substation. 



336 



with the necessary self-contained ability to detect line over- 
currents, to time and interrupt the overcurrents, to reclose 
automatically, and to reenergize the line. If the line overcur- 
rent is permanent, the recloser will lock open after a preset 
number of operations (usually three or four) and isolate the 
failure. Thus, the recloser can eliminate prolonged outages 
of the distribution system due to temporary faults or tran- 
sient overvoltage conditions (12). 

Reclosers can be hydraulically or electronically con- 
trolled (12). With the hydraulic control, an overcurrent is 
sensed by a trip coil in series with the line. When the 
minimum trip current of the recloser is exceeded, the trip 
coil actuates a plunger and causes the recloser contacts to 
open. The timing and sequencing are accomplished by the 
pumping of oil through separate hydraulic chambers. 
Electronically controlled reclosers provide a more easily 
adjusted, flexible, and accurate control then the hydrauli- 
cally controlled recloser. The electronic control gives a 
convenient means for changing time-current characteris- 
tics, trip-current level, and the sequence of the recloser 
operation without deenergizing the recloser. Auxiliary 
tripping devices are available to allow additional protec- 
tion, such as ground-fault and ground-check monitoring. 
Activation of the auxiliary tripping causes the recloser to 
lock open. 

The selection and application of a recloser is basically 
the same as for an interrupter. The necessary items to 
consider are system voltage, maximum available fault 
current at the recloser location, maximum load current, 
maximum fault current in the zone protected by the 
recloser, coordination with other protective devices up- 
stream and downstream, and ground-fault tripping or 
sensing. The common ratings are voltage, basic impulse 
insulation level (BID, continuous current, minimum trip 
current, and interrupting current. 

Disconnect Switches and Fuses 

To provide a visual disconnect for maintenance pur- 
poses, knife-blade load-break switches or fuse cutouts are 
located on the primary and secondary of the substation 
transformer. The switches are usually housed in metal- 
clad enclosures, while the cutouts are pole mounted. 
Gang-operated switches are recommended over hook-stick 
devices, since pulling one line at a time with the circuit 
under load can obviously single-phase the system. 

Fuse cutouts have an interrupting capacity instead of 
a short-time current rating. They should be used to isolate 
parts of a deenergized circuit, even if the cutouts used are 
designed with mountings that operate as a load-break 
switch. Cutouts alone cannot be relied upon to protect the 
transformer in all cases, especially when the available 
short-circuit current is above their interrupting rating. 
Power fuses should therefore be used as backup protection. 

High-voltage fuses may be employed as the main 
means of transformer protection. As with any protective 
device, their time-current characteristics must be coordi- 
nated with upstream and downstream devices, here the 
utility and the protective relaying on the secondary. Fuses 
alone cannot provide ground-fault protection of the trans- 
former primary. When fuses are used and ground-fault 
protection is needed, relaying could be used to trip the 
mechanism of an automatic load-break switch. A general 
protection rule, though, is to use a primary circuit breaker 
with protective relaying when the three-phase trans- 
former capacity is 5,000 kVA and above (13). For any 



capacity, the primary breaker prevents the single phasing 
that fuses allow and permits easier coordination with 
other relaying. 



PROTECTIVE RELAYING IN SUBSTATIONS 

The sizing of overload, short-circuit, and ground-fault 
relays is basically no different from that described already 
in this and previous chapters for other power equipment. 
Transformer protection is a protective-relaying problem in 
substations. With larger substation capacities, relay 
pickup for overloads and short circuits alone is normally 
too high to provide adequate protection. Sudden-pressure 
relays that sense the rate of pressure rise in the gas 
cushion of the transformer tank are sensitive to small arcs 
under oil, and their use is warranted (13). Further, 
percentage-differential relaying is also strongly suggested 
for internal fault protection of transformers rated at 5,000 
kVA or higher, although it is not as yet a widespread 
practice. 

A standard percentage-differential relaying system is 
illustrated in figure 13.16. Figure 13.17 is a typical 
one-line diagram of a substation with this differential 
relaying added. The CT's on a wye-connected winding of a 
transformer should be connected in delta, while CT's on a 
delta winding are connected in wye. There are two basic 
requirements that percentage-differential relaying must 
satisfy (11). 

1. The relays must not operate for load currents or 
external faults. 

2. The relays must operate when internal faults are 
severe enough. 

Relay pickup is used to trip the breaker on the transformer 
primary. 




Operating 
coil 

Restraining 
coil 



Percentage - 

difference 

relay 



Figure 13.16.— Standard percentage-differential relaying 
system for transformer protection. 



H^M^H^ 



337 



Utility 



59-G 




Operating 
coil 



Restraining 
coil 



— > 

To borehole 



Ground 



Power 



Pilot 

Figure 13.17.— One-line diagram of substation with 
percentage-differential relaying. 



Differential relays usually have coil taps to compen- 
sate when the CT's are not perfectly matched. When 
selecting a CT for differential relaying, the common prac- 
tice is to choose the highest CT ratio that will provide a 
secondary current as close as possible to the lowest-rated 
relay tap. This minimizes the effect of impedance in the 
wiring connections between the CT's and the relays. To 
assure that the relay will operate a maximum sensitivity, 
the current supplied to the relay under maximum load 
conditions should be as close as possible to the continuous- 
current rating of the tap (11). 

Percentage-differential relays usually have an adjust- 
ment to vary the percent slopes. The adjustment provides 
a means for preventing unreliable relay operation due to 
unbalances between CT's during external faults. Unbal- 
ances can occur from the following: 

• Tap changing of the power transformer, 

• Mismatch between CT secondary currents, and 

• The difference between CT errors on either side of 
the power transformer. 

If a power transformer is rated at 10,000 kVA or 
greater, a harmonic-restraint circuit is recommended in 
addition to the percentage-differential relays (13). This 
circuit causes a differential relay to be self-desensitizing 
during magnetizing inrush periods, but the relay is not 
desensitized if a short circuit occurs in the transformer 
during a magnetizing inrush period. Only the fundamen- 
tal component of the differential current is delivered to the 
operating coil. The harmonics are separated, rectified, and 
delivered to a restraining coil (11). 



The pressure relays mentioned earlier can play a 
valuable role as supplemental protection to differential 
relaying. In fact, with sensitive and reliable pressure 
relays, the sensitivity of the differential relays can be 
reduced to prevent undesirable operation due to inrush 
current (11, 13). Thermal-overload protection should also 
be provided as a third means of transformer protection. 



LIGHTNING AND SURGE PROTECTION IN 
SUBSTATIONS 

Since much of the equipment of the surface substation 
may be exposed, there must be protection against tran- 
sient overvoltages due to lightning as well as due to 
switching. Overhead static wires and shielding masts (see 
chapter 11) are commonly used to protect substation 
equipment from direct lightning strokes. Two static wires 
can be positioned above and between the line conductors to 
provide shielding for overhead distribution lines. In addi- 
tion, surge arresters are mandatory to limit the transient 
overvoltages to safe levels. 

The surge arresters on the incoming lines should be 
located as close as possible to the transformer terminals. 
Station-class valve arresters should be used. Again, ar- 
rester selection should be based on the primary voltage, 
the effectiveness of grounding, and insulation coordina- 
tion between the arrester and the transformer BIL. With 
dry-type transformers, the BIL is practically constant with 
the width of applied impulse (see chapter 12, figure 12.5). 
The margin of protection could then be less for the arrester 
sparkover voltage than for the IR-discharge voltage. It can 
be seen in figure 13.18 that in liquid-immersed transform- 
ers, the insulation withstand is not a linear function with 
the impulse width. Instead, the insulation-withstand level 
decreases from the front-of-wave to chopped-wave to the 
full-wave values (12). The full-wave value is the BIL 
rating. Standard values for oil-immersed transformers are 
listed in table 13.2 (9), and values of insulation classes 
below the standard are provided in chapter 11, table 11.1. 
As shown in figure 13.18, the BIL should be compared 
with the discharge voltage. Because the exposure avail- 
able in substations can lead to worst case surge conditions, 



Table 13.2.- 


-Standard BIL's for oil- 
transformers 


immersed power 




Primary winding 
phase-to-phase voltage, 


V 


Insulation 
class, kV 


BIL, 
kV 


22,900... 








25.0 
25.0 
34.5 
34.5 
46.0 
46.0 
69.0 
69.0 
92.0 
115.0 

138.0 

161.0 


150 


23,000... 








150 


26,400... 








200 


34,500... 








200 


43,800... 








250 


46,000... 








250 


67,000... 








350 


69,000... 








350 


92,000... 








450 


1 1 5,000 . 








550 


138,000. 








1 350 

M50 

650 


161,000. 








1 550 

1 450 

750 










'650 
1 550 



1 Reduced BIL's may be applied if proper coordination is maintained with 
surge arresters. 



338 



> 
< 



o 
> 



Front-of-wave 
withstand 




Chopped-wave 
withstand 



BIL (full-wave withstand) 



Front-of-wave 
sparkover 



Minimum margin 
* of protection 

^-Discharge-voltage maximum 
r characteristics of arrester 




Impulse voltage wave rising 100 kv///s 
for each 12 kV of arrester rating 



1 



3 4 
TIME, //s 



Figure 13.18.— Insulation characteristic of liquid-immersed 
transformer compared with the characteristic of valve surge 
arrester. 




ii f 

(i 1> 1 

ii < ,, 

<i 



Safety 

ground 

bed 



I = 25ft or 2d, 
whichever is larger 



Substation ground 
mat 

Figure 13.19.— Plan view showing locations of system and 
safety ground beds. 



the use of 20,000-A discharge current to establish the 
IR-discharge voltage is probably justified. The margin of 
protection should not be less than 0.20. 

Surge arresters on the outgoing lines should be lo- 
cated as close to the point they leave the substation as 
practical. These can be station-class arresters but are 
commonly intermediate class. 

Since higher insulation levels are common on break- 
ers and switches, at more remote locations it is sufficient 
to provide protection with arresters only. Breakers and 
disconnect switches are designed so that the impulse- 
withstand level over open switches is greater than that to 
ground, which reduces the possibility of flashover. As a 
result, a surge on these devices is likely to flash to ground 
and will not usually cause permanent damage (13). 

Whenever current-limiting fuses are applied with 
surge arresters, the possibility exists that the arc voltage 
produced by fuse operation could result in arrester spark- 
over and damage (13). Station-class and intermediate-class 
arresters are more susceptible to damage, because of their 
lower protective characteristics. If fuses are located on the 
line side of the arresters, arrester sparkover will not occur. 



SUBSTATION GROUNDING 

Extensive discussion of grounding, including informa- 
tion about substation applications, is presented in chapter 
7, but there are specific concerns that could not be covered 
until now. Before discussing them, a short review should 
be helpful. 

Two separate ground beds are required at the substa- 
tion: the system bed and the station bed. The station bed 
is usually a mesh located underneath the substation area 
and is sometimes called the substation ground mat. As 
shown in figure 13.19, the safety bed is located at a 



distance from the station ground bed (3). Surge arrester 
grounding conductors, static conductors, metallic frames, 
and fencing within the substation area are connected to 
the station ground bed. Only the distribution grounding 
conductors, including that of the grounding resistor, are 
tied to the safety ground bed. 

Separation is mandated to prevent the voltages that 
might be produced across the station bed from being 
transferred to the safety bed. This is particularly critical 
when portable or mobile equipment is involved, since all 
equipment frames can be assumed to be at the potential of 
the safety ground bed (that is, the voltage drop across the 
bed to infinite earth). By law, the two beds must be 
separated by a minimum of 25 ft. However, better separa- 
tion is achieved by a distance not less than two times the 
radial influence of the safety ground bed. This radial 
influence is defined as the longest straight line that can be 
drawn within the volume enclosed by the safety bed (3). 
The separation distance is measured between the two 
closest points of the system and the safety beds. To help 
minimize possible coupling between the two ground beds, 
the safety bed should be placed in a location adjacent to 
the corner or narrow side of the substation (fig. 13.19). 

To limit the magnitude of voltage produced across a 
bed during current flow, low resistance is mandatory for 
both ground beds. The resistance value is generally de- 
fined as 5 fl or less. To ensure that step and touch poten- 
tials are not hazardous in the grounding location, poten- 
tial gradients during surging conditions must also be 
restricted. The design of an adequate safety ground bed 
that will meet these requirements is presented in chapter 
8. The guidelines for the substation system ground bed are 
covered in the following section; this information has been 
abstracted from Bureau of Mines Information Circular 
8835, an excellent reference on the subject (3). 



339 



Substation Ground Mat 

The system ground bed is a series of interconnected 
conductors and possibly ground rods buried in the earth 
under the substation. The main purpose of this ground bed 
is to provide electric-shock protection for personnel in and 
around the substation during lightning strokes, short 
circuits, equipment failures, and many situations involv- 
ing human error. This protection is provided by the mul- 
tiple grounding of all accessible surfaces within the sub- 
station area, to limit step and touch potentials to safe 
levels. Substation equipment that cannot be grounded is 
placed so it is inaccessible and cannot be touched by 
personnel. The second purpose of the system ground bed is 
to limit insulation stress by conducting surges to earth. 

One of the most important aspects of the substation 
design is the construction of the ground bed itself. The 
purpose of the system ground bed is to hold the voltage of 
the substation floor at ground potential. A common hazard 
within the substation occurs from touch potentials that 
are too high; these potentials can be held to acceptable 
levels by using an open grid of conductors buried beneath 
the surface. The maximum acceptable grid spacing de- 
pends on the available fault current, resistivity of the 
earth, clearing time of protective devices, and overall bed 
geometry. Regardless, the following guidelines should 
provide adequate safety in practically all situations (3, 6). 

The substation should be enclosed completely by a 
continuous buried ground conductor. This perimeter con- 
ductor should extend approximately 3 ft beyond the outer 
substation fence to protect any person who might contact 
the fence. Within the perimeter conductor, a regular grid 
or mesh of wires spaced approximately 5 to 15 ft apart 
should cover the entire substation area. These wires 
should be uniformly spaced and located along rows of 
equipment to facilitate grounding connections. Extra con- 
ductors should be added to the grid at the corners of the 
substation and in substation work areas. 

The conductors should be bonded together at all 
intersecting points of the mesh, using either exothermic- 
welded connections or heavy clamps designed for ground- 
ing applications. All substation equipment should then be 
bonded to the ground mesh at two different points, prefer- 
ably at points of interconnection on the mesh. The substa- 
tion fence, posts, and gates should also be bonded to the 
grounding grid. The substation site should be chosen to 
avoid gas pipes, water pipes and other buried conductors if 
at all possible; otherwise, these ungrounded conductors 
must also be bonded to the grid at several points. 

Borehole casings require special treatment because 
they can provide a low-resistance path from surface areas 
to underground workings (3). The boreholes within the 
substation should be bonded to the station grounding grid 
to protect personnel in the substation area. Protection 
cannot be provided to both ends of the casing; therefore, 
the casing extending underground must be made inacces- 
sible for at least a 10-ft radius and labeled as dangerous. It 
must not be used as ground reference for any equipment or 
haulage system underground. If the borehole casing is 
outside the substation but in the vicinity of the safety 
ground bed (fig. 3.19), it should be coupled to the safety 
ground. In this case, the underground end of the borehole 
casing can be used as a grounding reference. The borehole 
supporting structure should also be bonded to the casing 
but isolated from the station ground so that surface 
personnel cannot contact both grounds simultaneously. If 



the borehole is outside the substation and away from the 
safety ground, it can be left ungrounded. Here, both ends 
should be avoided but isolation is not mandatory. (For 
additional discussion, see reference 3, page 23.) 

If a low resistance is difficult to obtain or if the area 
experiences severe freezing or drying, ground rods can be 
driven at mesh points and interconnected with the ground- 
ing grid. These rods are most effective if they are driven 
near the periphery of the bed, especially near the corners. 
However, driving the rods closer together than 10 ft has 
not been found effective in lowering ground-bed resistance. 

All buried conductors must be corrosion resistant. Cop- 
per should never be mixed with steel or aluminum; for 
example, if steel clamps are used with copper conductors, 
they will be rapidly destroyed by corrosion unless protected. 

To provide adequate mechanical strength, the size of 
the mesh conductors should be not less than 2/0. Prefab- 
ricated meshes are available that use numerous small 
conductors welded together; this is also mechanically 
acceptable. The conductors can be copper, copper-clad 
steel, galvanized steel, or stainless steel, but it is usually 
not a good idea to use aluminum unless an alloy is 
available. 

The ground-bed conductors should be buried about 18 
to 24 in below the substation grade. The perimeter con- 
ductor can be placed somewhat deeper to reduce the 
potential gradients. Closely spaced conductors near equip- 
ment manual control handles should be installed closer to 
the surface for maximum protection. Rubber mats or 
wooden platforms can be used for temporary personnel 
protection at control handles instead of closely spaced 
conductors, but these mats must not remain in the substa- 
tion permanently since they are usually not effective when 
wet and can be a hazard in the winter. Special care must 
be taken to ensure that the insulated area is large enough 
to protect the area around the handle. 

The use of a gravel surface in the substation greatly 
improves safety because it significantly increases contact 
resistance, thereby increasing tolerable touch potentials 
by a factor of 2 or more. The gravel should be well drained, 
have a thickness of at least 4 in, and extend no less than 4 
ft beyond the substation fence. 

The total resistance of the system ground bed is 
closely related to its perimeter, and a long narrow bed will 
have a lower resistance than a square bed of the same 
area. It should be noted that the total resistance cannot be 
lowered by decreasing the grid spacing nor by driving 
ground rods within the substation area. Thus the mat 
should be designed to cover a relatively large area with 
conductors arranged to reduce high-voltage gradients. 
Although the equipotential characteristics of the system 
ground bed will protect personnel regardless of its resis- 
tance to earth, it is mandatory that this resistance be low, 
5 J2 or less. This limits the voltage rise during lightning 
discharges, thus limiting the stress on power-system insu- 
lation. A lower ground-bed resistance also tends to reduce 
potential gradients in the earth around the substation. For 
moderate to large substations, the required size of the 
system ground bed is usually big enough to ensure that 
the resistance is low. However, it may be necessary to 
augment the earth connections with driven rods or addi- 
tional conductors in the case of small substations or 
substations constructed on high-resistivity soil. Figures 
13.20 and 13.21 illustrate suggested system ground beds 
for large and small substations (3). 



340 



Several additional comments about substations that 
feed underground mines are in order. First, it is recom- 
mended that the safety ground bed be located as close as 
practicable to the borehole or the point where the power 
extends underground. This practice minimizes the possi- 



Fence 



JM: 



It 



\ 



Operating 
handle 



Borehole 
casing 



'~1 — 1 — VT 

... 1 1 IX. 



Main work area 



~7\ 

-t- 



^ 



10 



Exit area (wide enough 
to contain open gates) 



Scale, ft 



PLAN VIEW 




ELEVATION 

Figure 13.20.— Typical system ground bed for large substa- 
tion. 



Fence 




Exit area 
(wide enough 

to contain 
open gates) 



PLAN VIEW 




Open 
gate 



k 



Grade level 



ELEVATION 

Figure 13.21.— Typical system ground bed for small substa- 
tion. 



bility of damage to high-voltage circuits entering or ex- 
tending underground, which can lead to hazardous condi- 
tions, and reduces the probability of lightning strokes 
hitting exposed high-voltage circuits extending under- 
ground. To further reduce these problems, electrical cir- 
cuits should have the following protection within 100 ft of 
the point on the surface at which they enter an under- 
ground mine: 

• Overload and short-circuit protection, 

• Ground-fault protection, 

• Surge protection, and 

• A visual disconnect. 

As all of these can be contained within the substation 
area, it is logical to place the entire substation as close as 
possible to the underground power entry. This can also 
prevent unnecessary duplication of components. 

Ground-Fault Protection 

A resistance-grounded system is mandatory for porta- 
ble and mobile equipment being fed by the substation. The 
ohmic value of the grounding resistor must be such that 
the frame potentials of the high-voltage distribution are 
no greater than 100 V during ground-fault conditions. For 
practical purposes, this restricts the maximum ground- 
fault current to not more than 50 A. Most modern substa- 
tions serving mines use a 25-A ground current limit, 
however it is important that the selected current limit is 
greater than three times the per-phase capacitive- 
charging current. 

Ground-fault protection at the surface substation 
should be by 

• Zero-sequence relaying for each outgoing circuit, 

• Potential relaying about the grounding resistor, and 

• (Direct) neutral relaying between the grounding 
resistor and the neutral bushing of the power transformer 
or the grounding transformer. 

Protective relaying to detect grounding-transformer fail- 
ures is wise, but protection such as fusing should never be 
used if it will remove the transformer from the line or the 
grounding resistor. Ground-check monitoring of circuits 
feeding portable and mobile equipment is essential, and it 
may also be necessary on circuits feeding other loads. 



ADDITIONAL MINE SUBSTATION LOADS 

When other loads are being supplied from a mine 
substation, such as fans and preparation plants, ground- 
ing problems can become complex. Three fundamental 
concepts must still be served: 

1. Provision of ground-fault protection, 

2. The earth cannot be used as a grounding conduc- 
tor, and 

3. Isolation of the grounding system serving portable 
and mobile equipment. 

Adherence to these criteria is mandatory where under- 
ground mines are connected with the substation. The 
following discussion will be pointed directly at these 
applications, although the comments are adaptable for 



341 



any mine. Figure 13.22, illustrating a substation feeding 
both surface and underground loads, will be used as a focal 
point. 

At first glance, the circuits shown in this figure would 
appear to be reasonably safe. The actual surface load is 
isolated from the main substation secondary through the 
transformer windings of its unit substation. Both substa- 
tions have separated ground beds with all appropriate 
connections. Each outgoing circuit from the main substa- 
tion is separately protected for line and ground overcur- 
rents. However, consider the situation where a ground 
occurs in the unit substation, for instance in the incoming 
circuitry or within the transformer. This can establish a 
ground current path through the earth, the main substa- 
tion safety bed, and its grounding resistor. For the time 
that earth current flows, the mine grounding system is 
elevated by a potential equal to the mine ground-bed 
resistance times that current. In other words, one forbid- 
den grounding practice is shown in the figure: the earth is 
being used as a grounding conductor. Furthermore, the 
ground-fault protection used in this system portion, zero- 
sequence (51G) and potential (59G) relaying, must have a 
specific level of ground current to pick up. Earth resistivity 
may be large enough that detection is not made and the 
current persists. Even if tests are made to ensure pickup, 
seasonal variations might destroy all intentions. 



Figure 13.23 illustrates a solution to this problem. 
Here, the frames of all direct loads for the main substation 
are connected via grounding conductors to the safety 
ground bed. Ground-check monitoring of these conductors 
to the surface facilities is preferred, but equal safety may 
be obtained by establishing a low-resistance ground bed on 
the unit substation end. The new bed must be considered 
as part of the mine ground bed. 

Because of the increased number of components con- 
nected to the safety bed, the probability of lightning 
strokes and other surging conditions elevating the bed is 
also raised. Surge protection must therefore be added. 
Static lines should protect overhead power, grounding, and 
ground-check lines extending to the surface loads. The 
static lines must not be connected to the mine ground bed. 
Outgoing-circuit grounding of the unit substation should 
be separated from the mine ground bed by the same 
provisions as station ground beds. 

The recommendations above do have exceptions. First, 
if the transformer used for the surface loads is in the same 
substation area as the main power transformer, a ground- 
ing conductor to the mine ground bed must not be used, 
because this could easily tie the substation system and 
safety beds together. In some instances, isolation of the 
unit substation secondary might not be practical or could 
create a personnel hazard. An example is a unit substation 



SURFACE FACILITIES 
UNIT SUBSTATION 



MAIN SUBSTATION 



50/51 51-6 

.... o— , 



Or other acceptable 
practice 



Fencing 




Power 



3»- Grounding 



V 

Separation of 
system and 
mine beds 



Figure 13.22.— Substation feeding both surface and underground loads (no grounding conductor). 



342 



Static protection 
of surface lines 



SURFACE FACILITIES 
UNIT SUBSTATION 




Shall not be 
connected to mine 
ground (if isolated 
delta, ground system 
shall have separate 
grounding electrode) 



Figure 13.23.— Substation feeding both surface and underground loads. 



mounted on a metallic structure associated with a prepa- 
ration plant. An attempt at isolation here could result in 
two proximate surfaces that could have dangerous poten- 
tials existing between them. In this situation, the ground- 
ing conductor to the mine bed is still mandatory, yet 
exposure of the mine bed to lightning-stroke surges would 
be substantially increased, without the next exception. 

It can be shown that if a grounding conductor is 
approximately 250 ft in length, its typical impedance to a 
surge (such as an 1.2 x 50-/*s impulse) is high enough to 
stop the traveling wave effectively. Thus, if a safety factor 
of 2 is applied, any grounding conductor of 500 ft or longer, 
protected by a static line or otherwise, can be connected to 
the mine ground. This means that when isolation of the 
unit substation is dangerous, the mandatory grounding 
conductor to the mine bed should be effectively 500 ft long. 
In addition, if a grounding conductor is protected by static 
wire for 500 ft, it can be used for shielding the power 
conductor provided that it is properly grounded beyond 
that distance. 

When these precautions are met, the result should be 
a safe substation grounding system, and yet there is still a 
high chance that personnel safety could be compromised 
through human error. Hence, maximum safety will be 
achieved if a separate substation is used to supply mine 
power. Perhaps the best but most expensive approach is to 
use subtransmission to feed a second substation for sur- 
face facilities, but use of an isolation transformer between 
the main substation and the mine provides the same 
benefits. This will be covered subsequently. 



PORTABLE SUBSTATIONS 

As mentioned earlier, portable substations have 
gained recent popularity for small load situations in 
mines. The entire substation is skid mounted and can be 
transported as a unit by truck. Portable substations are 
custom built to meet the needs of the individual mine. For 
some small mines, the economics are such that the porta- 
ble substation is used in conjunction with a diesel-powered 
generator. The output of the generator may be a low 
voltage, such as 480 V, and here the substation trans- 
former may step up the voltage to 4,160 V for distribution 
to the mine loads. 

Figure 13.24 shows a situation where the utility 
company supplies a voltage of 12.47 kV. This particular 
substation is designed to operate a mine load of 750 kVA 
and surface loads of 150 kVA, including a conveyor belt 
and a ventilation fan. The surface and the mine circuits 
are in parallel with each other and supplied by the 
incoming feeder. The surface circuit can be isolated from 
the mine circuit by means of the load-break switch, and 
the primaries of both the surface and mine circuits are 
protected by high-voltage fuses. The surface transformer 
steps down the incoming line voltage to 480 V for operat- 
ing the belt and fan. Each of the surface loads is protected 
by a molded-case circuit breaker that contains an under- 
voltage release powered from the utility-control trans- 
former. The utility-control transformer also supplies power 
to the 240- and 120-V outlets. 



343 



Molded- 

case 
breakers 



12,470 A -480Y/277 
150 kVA 





12,470 A-4.160Y/2400 
750 kVA 




To 51-G 



To belt 



To fan 



Spare 




^ To mine 
loads 



I 

To 50/51 



CT - *" To control 
p circuit 

4,160-120 



Figure 13.24.— Typical portable substation to service small mine. 



The mine transformer steps down the incoming volt- 
age to 4,160 V for distribution to the mine loads. The 
transformer secondary is wye-connected to permit resis- 
tance grounding and feeds the VCB, which in turn pro- 
tects the distribution circuit. CT's at the breaker output 
supply inverse-time relays for line-overcurrent protection 
and ground-fault protection. The control transformer on 
the line side of the VCB supplies power for tripping the 
breaker. 

As in larger permanent surface substations, individ- 
ual ground beds must be provided for the station ground 
and the safety ground. These connections are noted by the 
two different ground symbols in the schematic. 



UTILITY VOLTAGE AS MINE DISTRIBUTION 

Some mines must purchase power at the same voltage 
as mine distribution, either because of economics or as the 
result of a utility company decision. The utility company 
often supplies power from an ungrounded delta trans- 
former secondary. The mine operator must supply basi- 
cally the same substation components as those discussed, 
but the main transformer may or may not be included, 



depending on the location of the utility transformer (which 
is the power source or source transformer) and also on the 
specific location where a neutral must be derived to 
establish mine grounding. For surface mines, this location 
is not always critical, but for underground mines the mine 
ground bed should be as close as practical to the point 
where circuits extend underground. Regardless, the follow- 
ing rules apply to all ungrounded delta sources supplying 
portable or mobile mining equipment, including situa- 
tions where the mine owns the source transformer. 

If the distance between the source transformer and 
the grounding location is 100 ft or less, the simple ar- 
rangement in figure 13.25 can be used. A grounding 
transformer derives the neutral point, a grounding resis- 
tor is inserted to limit ground-fault current, and a safety 
ground bed is established. Load-break switches provide 
visual means of disconnect, and surge protection is af- 
forded by arresters grounded to an arrester ground bed 
separated from the mine ground. Protective relaying is 
identical to that in any substation: line-overcurrent, 
ground-fault, and ground-check monitoring. Here it is 
imperative that the recommended zero-sequence relaying 
be located downstream from the grounding-transformer 
connection, otherwise it cannot detect downstream ground 



344 



faults. Zero-sequence currents terminate at the grounding- 
transformer terminals. From switch to switch, this addi- 
tional equipment is defined as the substation for the mine. 

When the distance is greater than 100 ft, the one-line 
diagram of figure 13.26 applies. This arrangement re- 
quires the installation of an isolation transformer, but 
otherwise the required substation is practically identical 
to that in figure 13.11. Although a delta-wye transformer 
is shown, a delta-delta power transformer with a ground- 
ing transformer could also be used. This system reflects 
the Federal regulation that requires the grounding resis- 
tor to be located at the source transformer. To be consistent 
with other regulations, "at" is defined as 100 ft or less, 
and the logic for the requirement is very sound. 

Utility companies commonly supply the ungrounded 
delta power on overhead lines. The typical distance for one 
overhead span (pole-to-pole) is approximately 100 ft. Ex- 
posed power lines, beyond one span, do raise the probabil- 
ity of lightning striking the mine power system, but 



equally important, the possibility of losing a line conduc- 
tor is significantly increased. One common example is a 
vandal's shooting an insulator or power conductor with a 
rifle or handgun. If one of the lines between the source 
transformer and the grounding transformer is lost (for 
example, in figure 13.25), the grounding transformer can- 
not find the system neutral, hence its common terminal, 
which is connected to the grounding resistor, will rise to 
almost the same potential as a remaining line conductor. 
With normal system operation, the problem presented 
here is minor since the resulting single phasing would be 
quickly detected by the overcurrent protective circuitry. 
Under no-load or low-load conditions, however, this cir- 
cuitry might not detect the unbalance, and all potentials 
in the mine could become elevated to near line potential. 
Potential relays across each leg of the grounding trans- 
former can be employed to detect such a hazard and supply 
tripping to the main circuit breaker. An isolation power 
transformer provides this same protection. 



100 ft 
or less 



Control 
power 



50/51 51-G 



ii 

^ 



Utility I f 

company Zig-zag 

transformer transformer 



Ground [ 
resistor 




Substation fencing 



"5 



c—\ j Ground- 

X Wop check 
c — i 3»"^-circuitry 

' (37) 



Pilot | I 



Grounding j 



System 
(substation) i Mine (safety) 
ground bed ' ground bed 

« 25 ft — *| 



Figure 13.25.— Providing mine ground and protective relay- 
ing from utility substation. 



If fed underground, must 
be shielded construction. 
For surface, may be overhead 
lines or shielded cable. 




company 
transformer 



Substation fencing 
■\j Syste 



m 
(substation 
ground bed 

> 25 ft 



Figure 13.26.— Use of isolation transformer with utility 
substation. 



ADDITIONAL SUBSTATION DESIGN 
CONSIDERATIONS 

Over the years many additional design precautions 
have been developed through experience with permanent 
substations at mines. A listing of some of these follows: 

• Substation conductor insulators can be of post or 
cap-and-pin types, and either standard strength or heavy- 
duty design can be used (13). "Extra-creepage" insulators 
and bushings (or one voltage class higher than required) 
are recommended (7) as they extend the time between 
required cleanup and minimize flashover. 

• The high-voltage side should be located on the 
windward side of the substation (9-10). The prevailing 
winds will then help keep the substation insulators clean 
between regular cleanups. 

• The voltage drop on the distribution system should 
be maintained within + 5% to all load concentrations (7). 
This establishes the nominal voltage and necessary taps 
for the secondary of the substation, and corresponds to the 
taps stated earlier for transformer primaries in power 
centers. 

• Permanent substations should be located outside 
the influence of the mine. This is very critical for surface 
mines. 

• Buses and bus supports should be designed to 
withstand the stress from the average asymmetrical short- 
circuit current during the first 10 cycles after a fault (13). 

• Line tensions within the substation area should be 
assumed to occur under maximum wind-loading and ice- 
loading conditions (13). Structure design should also be 
based on these specific line tensions and wind loading. 

• Substation primary voltages above 34.5 kV cannot 
be used for distribution-class equipment (that is, cannot be 
transformed directly to low-voltage and medium-voltage 
mine usage) (13). This could create problems in supplying 
certain surface loads. 

• The utility will meter the substation power con- 
sumption on the primary or secondary, depending on the 
primary voltage. The substation should also contain the 
mine's own metering, of equivalent precision to that of the 
utility (7). This is invaluable for maintenance checks and 
also serves as a double-check on the utility. Metering could 
include kilowatt, ammeter, voltmeter, and power-factor 



345 



instrumentation for each outgoing feeder from the substa- 
tion. Demand meters, for either 15- or 30-min maximum 
demand, may also be needed. 

• The fence enclosing the substation area should be 
no less than 8 ft high (1). The fence should be provided 
with a door or gate, which should be locked except when 
authorized personnel are present. The fence should be 
grounded to the station ground bed. It should also extend 
to the ground so no one can crawl underneath. Danger 
signs should be posted on all sides of the substation area. 

• The substation area should be maintained free of 
weeds, trash, and combustible material that might create 
a hazard to personnel or the substation components (2). 

• All components containing liquid insulation should 
be mounted on pads to allow quick drainage and capture of 
any lost fluids. This is especially important when flamma- 
ble or environmentally hazardous liquids are involved. 

Chapters 12 and 13 have attempted to collate an 
enormous amount of information to provide a coherent 
presentation of typical construction practices used for 
mine power equipment. Some designs, like mine power 
centers, are so dependent upon the individual application 
that standard units cannot exist. At the same time, the 
assembly of others, such as switchhouses and substations, 
is rather uniform across the industry. Because the subject 
matter is so large, some generalization has been essential 
and the omission of some specialized material was inevi- 
table. Any engineer involved with mine equipment should 
of course work very closely with the manufacturer to 
determine the best solution to the specific needs of the 
mine. 



REFERENCES 

1. American Standards Association. American Standard Safety 
Rules for Installing and Using Electrical Equipment in and About 
Coal Mines (M2.1). BuMines IC 8227, 1964. 

2. Bifulco, J. M. How To Estimate Construction Costs of Elec- 
trical Power Substations. Construction Publ. Co., 

3. Cooley, W. L., and R. L. King. Guide to Substation Ground- 
ing and Bonding in Mine Power Systems. BuMines IC 8835, 1980. 

4. Dornetto, L. D. The Importance of Grounding Systems in the 
Protection of Personnel and Equipment. Paper in Mine Power 
Distribution. Proceedings: Bureau of Mines Technology Transfer 
Seminar, Pittsburgh, Pa., March 19, 1975. BuMines IC 8694, 1975. 

5. Institute for Electrical and Electronics Engineers (New 
York). General Requirements for Liquid-Immersed Distribution, 
Power, and Regulating Transformers. Stand. C57.12.00-1979. 

6. Guide for Safety in Substation Grounding. Stand. 

80-1976. 

7. Recommended Practice for Cement Plant Power 

Distribution. Stand. 277-1967. 

8. Recommended Practice for Electric Power Distribu- 
tion for Industrial Plants. Stand. 141-1986. 

9. Lordi, A. C. Electrification of Coal Cleaning Plants. 
Mechanization, v. 20, Oct. 1956. 

10. Trends in Open-Pit Mine Power Distribution. Coal 

Age, v. 66, Jan. 1961. 

11. Mason, C. R. The Art and Science of Protective Relaying. 
Wiley, 1956. 

12. McGraw-Edison Co., Power Systems Div. (Canonsburg, 
PA). Distribution-System Protection Manual. Bull. 71022, undated. 

13. Wade, E. C, and M. E. Kunsman. Practical Considerations 
for Selectiong and Coordinating Electrical Components in Outdoor 
Mining Substations. Paper in Conference Record -IAS 13th An- 
nual Meeting (Toronto, Ontario, Canada, Oct. 1978). IEEE, 1978. 



346 



CHAPTER 14.— SOLID-STATE CONTROL AND RELAYING 1 



Except for a short introduction to electronics in chap- 
ter 5, discussion of solid-state devices has been avoided up 
to this point. The reason is that their use is rather new to 
the mining industry. However, their impact has been 
substantial, and if one area of electrical engineering can 
be singled out as having the greatest probable influence 
on future mining operations, it is electronics. This chapter 
will primarily discuss two solid-state applications that are 
already important: motor control and protective relaying. 



MOTOR CONTROL 

Almost all activity and interest in the solid-state 
control of industrial motors have centered around the use 
of the silicon-controlled rectifier (SCR) or thyristor. A 
model and circuit symbol for this four-layer, three-junction 
device are shown in figure 14.1. The outer two layers act as 
a p-n junction and the inner layers serve as an element to 
control that junction. The device has three external termi- 
nals: anode, cathode, and gate. 

If the anode is positive with respect to the cathode 
(forward biased) and if the gate is reversed biased in 
reference to the cathode, there exists a balance of electri- 
cal charges in the four layers, and current flow in inhib- 
ited. This process is termed forward blocking, and the SCR 
exhibits high resistance in both directions. When the gate 
is forward biased with respect to the cathode, the gate 
current upsets the electrical charge balance, and anode- 
to-cathode current can flow. Once this conduction starts, 
the gate loses all control; in other words, the gate can turn 
the thyristor on but not off. Gate control (or turning the 
thyristor off) can only be achieved by reverse-biasing the 
gate-cathode circuit once more and reducing the anode- 
cathode current essentially to zero. 

The typical characteristic curve for a thyristor illus- 
trated in figure 14.2 details the phenomenon just de- 
scribed. Voltages and currents are given with respect to 
the cathode. When the anode potential is positive and the 
gate is at cutoff (negative potential), the characteristics 
are similar to those when the anode is negative. At a 
specific positive potential, called the breakover voltage, 
the SCR will turn itself on. Here, anode current increases 
considerably, and the voltage drop is reduced substantially 
across the device. The breakover point can be altered by 
varying the gate current, and this is why the thyristor is 
so valuable in power control applications. 

The common technique for initiating conduction is to 
apply a current pulse at the gate (12). 2 The pulse alone 
neutralizes the thyristor blocking action at the desired 
breakover voltage. The simplest thyristor application is in 
the half-wave rectifier circuit shown in figure 14.3. Break- 
over is determined by the gate current, and thus the 
average value of dc through the load can be controlled by 
the gate control pulse (i G ). The process of starting the 
thyristor is often referred to as firing, and the angular 



1 The author wishes to thank D. J. Tylavsky, assistant professor of 
electrical engineering, Arizona State University, who prepared the original 
material for the static protective relaying section of this chapter while he 
was a graduate student at The Pennsylvania State University. 

2 Italicized numbers in parentheses refer to items in the list of references 
at the end of this chapter. 



position or timing of each pulse with respect to the source 
waveform is called the firing angle (angle a). In order to 
control ac, the load must receive both sides of the ac 
waveform. Two single-phase circuits that can be used as ac 
voltage controllers are shown in figure 14.4. The bidirec- 
tional arrangement uses two thyristors back to back and 
gives a symmetrical output with appropriate firing-angle 
pulses for the two gates. The unidirectional circuit with 
one thyristor and one diode presents an asymmetrical 
waveform to the load. The dc component thus created is 
considered a major disadvantage because the source wave- 
form must be ideal for the circuit to be of practical value 
(9). 

Voltage control for three-phase loads employs three 
single-phase controllers as shown in figure 14.5. Although 
a bidirectional circuit is illustrated, unidirectional circuits 
can also be applied. The dc components are mostly can- 
celled with the three-phase waveform, but the unidirec- 
tional circuits introduce a higher harmonic content to the 
line currents (9). 





Gate section 
, * , 




Anode 


P 


N 


P 


N 


Cathode 













A Gate 
terminal 



-oGate 



Anode o 



-o Cathode 



Figure 14.1.— Model and circuit symbol for thyristor. 



J 



Avalanche 
breakdown 



ON 
characteristics 

w 

Breakover 
voltage 

~\~ V-±- 

OFF 

characteristics 



Figure 14.2.— Typical characteristic curve for thyristor. 



347 



'G 



VAK 



Q v=/2V sinort 



Vq = Vr CR 




' G An 



J2V| 
1.0 



IT 



tx 



1.0 

o- 




- cat 



2TT 



ojt 



■IT 27T 



-•►aft 



« IT 2TT 

Figure 14.3.— Thyristor half-wave rectifier. 



«rt 



Qi 
-K 1 



7*7^ 



Q. 



v=/2V sin art 



tt) 



Vo.'c 



R 



r\ 



+ . 



T+oc 21T 



oc 



U^ 



art 



(2.) 

Bidirectional 






v=/2V sin art 



Unidirectional 



v o^R 



v o>'o 


i 


(/.) 




R 
n 


T 


\ 






oc 


V 


/2TT * 



6Jt 



Figure 14.4.— Alternating current thyristor control. 



348 




Figure 14.5.— Three-phase control with bidirectional 
thyristor arrangement. 




Figure 14.6.— Full-wave thyristor bridge rectifier. 



Simple Motor Control 

One method of speed control in dc motors has already 
been covered, simply adjusting the dc voltage to the 
machine. This would apply to loaded series-wound motors, 
but shunt and compound dc motors can also be controlled 
by two direct means: armature voltage or field voltage. 
The field control is interesting because field excitation is 
usually considered constant, but in fact motor speed can be 
adjusted inversely with field current. The simple circuit in 
figure 14.3 could be used in any of these cases when the 
source is ac; however, the basic circuits of figures 14.6 and 
14.7 are commonly applied for single-phase and three- 
phase sources, respectively 

The commutation diode shown in figure 14.7 serves a 
specific purpose. Commutation is the transfer of current 
from one device to another as voltage relations change 
(21 ). The commutation diode conducts armature current so 
that it will not be transferred to the thyristors and diodes 
of the bridge and the thyristors can turn off at zero voltage. 
Without this diode, thyristor current might not reduce to 
zero and turnoff would not be likely. Other commutation 
techniques are available (9). 

A chopper or dc-to-dc converter is another method of 
variable-speed dc motor control using thyristors; it is 
especially important because it operates between a dc 
source and a load. As shown in figure 14.8, the thyristors 
in the chopper circuitry switch the dc source on and off, 
supplying unidirectional voltage pulses. This reduces the 
effective voltage, V , to the load. 

Variable-speed control of a standard three-phase 
squirrel-cage motor is possible using thyristor voltage 
control (for example, figure 14.5). Motor speed varies 
because the motor allows a greater percentage of slip at 
reduced voltage, then the controller increases the terminal 
voltage, the allowable slip is reduced, and the motor speed 
increases. However, three specific problems arise in this 
application. Reduced-voltage operation causes more heat 
to be generated in the motor. Motor torque is also reduced 
at slow speeds with lower voltage. Finally, the waveform 
produced by the thyristors is rich in harmonics (frequen- 
cies other than line) and this creates even more motor 
heat. Nevertheless, if high-quality motors are used and 
reduced-voltage operation is restricted to a short time, the 
result can be effectively controlled starting of squirrel- 
cage induction motors. 



Commutating 
diode 
\ 



s> c\ 



4 * 



- 6 



dc motor 
armature 

Figure 14.7.— Three-phase thyristor-controlled rectifier. 




v oA 
V 

Vr 







'ON 



tr 



'ON 

•» *■ 



'ON 



T — *U T— *l 

B 



Figure 14.8.— Simplified chopper control. 



349 



Control Systems 

lb this point, only thyristors and their application to 
variable voltage control have been discussed. Other than 
explaining how the device is fired, nothing has been stated 
about the control circuitry that supplies the gate signal. 
Perhaps the best way to visualize power control applica- 
tions with thyristors is through control-system principles. 
The two basic types of control systems are open loop and 
closed loop (12-13). In open-loop systems, the controlling 
element is unaware of the effect it produces in the con- 
trolled element, whereas in closed-loop or feedback sys- 
tems, information is delivered back to the controlling 
element so it can adjust control of the controlled element 
to produce a desired result. The feedback control systems 
of interest here are termed regulators whose purpose is to 
maintain the result constant. Feedback systems inher- 
ently provide more precise control than do open-loop 
systems. 

The operation of these systems can be described with 
the assistance of the basic block diagram in figure 14.9, 
where the lines in the diagram represent system variables. 
Each block represents a transfer function, the quantity 
that must be multiplied with the input to obtain the 
output. The comparison point in the block diagram is a 
position where two or more system variables are summed, 
and here the output is the algebraic sum of the entering 
variables. 

The variable to be controlled, C, may be voltage, 
speed, or any other quantity. The input or reference 
variable, R, is usually adjusted manually by an operator 
and determines the desired value of C. The feedback 
element, H, supplies information about the output to the 
comparison point, whose function is to compare the input 
and feedback signal. The difference or error, E, thus 
obtained drives the controller, G. In the simplest system, 
the controller might contain the device that needs to be 
controlled. (Note that G usually represents the transfer 
functions from input to output. H denotes transfer func- 
tions in a feedback path.) A feedback system contains all 
these elements and variables, whereas in open-loop sys- 
tems only the controller responds to the reference signal. 

Elaboration of these statements can be seen in the 
block diagram of figure 14.10, which illustrates the main 
parts of a motor controller (9). The power circuits, which 
primarily contain thyristors, provide variable direct or 
alternating voltage output to the controlled system. The 
controlled system can be a rotating machine or any other 
driven load. Appropriate feedback is supplied back to the 
controlling system and could, for example, represent cur- 
rent, voltage, temperature, speed, and so forth. The con- 
trolling system responds to the feedback and input infor- 
mation, and supplies control signals to the digital circuits. 
Finally, the digital circuits, acting as an additional trans- 
fer function, simply switch the thyristors in the power 
circuit off and on at appropriate times. 

Physical Characteristics of Thyristors 

The two common thyristor configurations used in 
industrial applications are illustrated in figure 14.11. 
Stud-mount types are intended for smaller loads and have 
average forward-current ratings up to around 150 A. Heat 
sinking is like that shown for diodes in chapter 5 and, 
because thyristors have p-n junctions, the same heat- 
dissipation theory applies to these devices also. The larger 



thyristor in the figure was introduced in the late 1960's 
and is termed a hockey puck, press pack, or disk (2). It has 
the ability to transfer junction heat on both sides and is 
clamped between two heat sinks as shown in figure 14.12. 





Comparison 


E 






t(f) 


Error or 

difference 

signal 








6 
Controller 




Input or 
reference 


-T 






Output or 


variable 






variable 
C 


R 












H 

Measurement 

and 

feedback 















Figure 14.9.— Basic control-system block diagram. 



Command 


Controlling 


->- 


Digital 
circuits 


-»- 


f 


D ower 






Controlled 


Response 




system 


circuits 


system 




















i 


Met 

Hit 
















Po 
inf 





Figure 14.10.— Simplified block diagram of a motor con- 
troller. 




Anode --^S 




Anode 



Stud mount Disk type 

Figure 14.11.— Common thyristor configurations. 




Figure 14.12.— Heat sinking of disk-type thyristors. 



350 



The heat-sink cooling can be by either air or water. Single 
disk thyristors presently have average forward-current 
ratings up to 7,500 A. Their relatively low cost has made 
solid-state motor control competitive in areas where prior 
to 1970 only electromechanical arrangements were consid- 
ered applicable. 

DIRECT CURRENT APPLICATIONS 

There has been excellent success in applying solid- 
state or static control to dc equipment, especially equip- 
ment with less than 100 connected horsepower (4, 17, 35, 
39). Examples include battery on-track and off-track vehi- 
cles as well as underground face equipment. The control 
has primarily used chopper circuitry, and two important 
advantages have been achieved: increased motor life by 
limiting armature current during acceleration, thereby 
reducing brush and commutator problems, and signifi- 
cantly reduced drive-train maintenance. General mechan- 
ical problems are reduced because motors have controlled 
acceleration. Even during plugging, which is the worst 
shock-loading instance, motors have controlled decelera- 
tion, reversal, then acceleration. For battery-powered ve- 
hicles, battery life is increased, since peak current demand 
is reduced and power is generally used more efficiently. 
(Applications in battery chargers are given in chapter 15.) 

Static control of dc motors is also applied in surface 
excavators, hoists, and conveyor belts, where there are 
variable-speed requirements and substantial demand fluc- 
tuation (39). The power supplied to surface excavators, for 



example, can swing from 200% demand to 120% regener- 
ation in 1 cycle. 

Currently, static control in surface excavators is lim- 
ited to machines of 20-yd 3 capacity or less. The form is 
similar to the conventional Ward-Leonard system (20, 39), 
but the motor-generator set is replaced by thyristors in a 
bridge configuration with feedback control. This change 
does improve operational performance, but it has a detri- 
mental effect on the ac distribution system to the machine 
(31). 

High-amplitude voltage transients and significant 
electrical noise commonly found on mine trolley systems 
have seriously hampered the adaptation of this technology 
to trolley locomotives (39). However, there has been some 
limited success with chopper control (4). It is interesting to 
note that problems have been encountered on practically 
all solid-state equipment receiving trolley power. 

An outstanding success of static control has been the 
ac-dc drive in underground ac face equipment. The design 
maximizes power-distribution efficiency, uses the traction 
advantages of dc series-wound motors, and has improved 
the performance available from thyristor control (35). 
Diagrams for systems used on shuttle cars and continuous 
miners are given in figures 14.13 and 14.14. In both, ac 
power is supplied through the trailing cable. After the 
main circuit breaker, ac is supplied through conventional 
contactors to the pump-conveyor (hydraulic) motor, the 
cutting motors, and an on-board three-phase transformer. 
The transformer secondary supplies ac for the dc systems, 
involving thyristor control of the traction motors. 



3- 

phase 

transformer 



ac 
switching 



Rectifier 



dc 
switching 



Control timing 
and 
feedback 



^ [ Traction \ 
~^~V motor J 



^ / Traction \ 
^^ motor I 



ac 

power 
switching 



3- 

phase 
power 
input 




Main 
circuit 
breaker 



Figure 14.13.— Block diagram of ac-dc shuttle car. 



351 



3- 

phase 

transformer 



ac 
switching 



Rectifier 



I 



dc 
switching 



Control timing 
and 
feedback 



( Traction \ 
"*"\ m otor J 

_^J Traction \ 
^^ motor J 



ac 

power 

switching 




Figure 14.14.— Block diagram of ac-dc continuous miner. 



ALTERNATING CURRENT APPLICATIONS 

Full variable speed cannot be accomplished with 
straight current or voltage control on ac induction motors. 
However, as motor speed is a function of the applied power 
frequency, speed control can be obtained by varying this 
frequency (12). A simple representation for one type of 
variable-frequency ac drive is shown in figure 14.15, and 
an elementary inverter circuit is given in figure 14.16. 
The incoming ac is rectified then filtered by capacitance to 
provide a high-quality dc. An inverter then converts the dc 
back to ac. By controlled switching of the thyristors, an 
alternating square-wave signal of the desired frequency 
can result. This can be used directly or filtered, as in the 
transformer in figure 14.16, to provide a sinusoid. Three- 
phase ac output can be obtained by employing three 
inverters and firing the inverters so that 120° timing is 
available. 

A direct adaptation of this technique has been used on 
production mining shovels as shown in the block diagram 
in figure 14.17 (20). This is reported to give better perfor- 
mance than Ward-Leonard dc motor control since machine 
characteristics are a function of the power electronics and 
not the motor. Another application of thyristor control of 
ac induction motors is in conveyor starters. 

Across-the-line starting of three-phase squirrel-cage 
induction motors on in-mine belt conveyors can be very 
detrimental. Belt conveyor installations can call for rather 
high horsepowers, and the resulting high starting current 



ac 
source 



-U 



Rectifier 



T 



Filter 



dc 

i 
I 
I 

i> 



Adjustable- period 

sinusoidal ac 

I 



Inverter 



V 



Filter 



ij 



Motor 



\T 



Rectified 


Adjustable-period 


Adjusted 


dc 


square wave 


speed 
output 



Figure 14.15.— Simple variable-frequency control. 




= v 



dc 



*- 






■Hr 



Figure 14.16.— Elementary inverter circuit. 



352 



On-board 
transformer 




Dynamic-braking 
control 



VFD 

INV 

RECT 

DYN BRK 

PM 

HM 

CM 

SM 



KEY 

Variable-frequency drive 

Inverter 

Rectifier 

Dynamic brake 

Propel motor 

Hoist motor 

Crowd motor 

Swing motor 



Figure 14.17.— Use of variable-frequency drive on production mining shovel. 



(three to eight times full load) can produce protective- 
relaying problems and large voltage drops. The latter can 
cause a motor torque decrease, which in turn can hamper 
belt conveyor acceleration. 

In the past, wound-rotor motors with step starters 
were used extensively on belt conveyors to overcome these 
problems. The step starter is essentially a bank of resis- 
tors that is connected to the rotor winding (see chapter 6) 
and allows the motor to start with a high resistance- 
to-reactance ratio for limited starting current and high 
starting torque. The external resistance is then decreased 
in steps, each decrease resulting in an increase in motor 
speed. At full speed, all external resistance is shorted out, 
and the wound-rotor motor operates like a low resistance- 
to-reactance design with corresponding high efficiency. 
Step starters for belt conveyor applications require large- 
capacity switching and contacting equipment. Histori- 
cally, these electromechanical components have been 
maintenance problems, and the brushes and slip rings of 
the motor can be a continual source of difficulty. 

On the other hand, controlled acceleration and lim- 
ited starting current can be achieved using squirrel-cage 
induction motors with solid-state starters. Compared with 
wound-rotor motors, motor and starter maintenance is 
lower, and the following advantages are gained: 

Reduced belt and splice tears; 

Decreased stress on mechanical power-transmission 
components, resulting in increased life; 

Elimination of some mechanical components, such as 
extensive gear trains, clutches, and so on; 

Decreased belt slippage, thereby reducing belt burns 
and removing impulse stresses; 



Custom design for special application; and 
Increased motor life. 

All these features have been validated by experience. 

Control Systems 

All three-phase solid-state belt starters in common 
use today employ reduced-voltage motor starting. The 
technique is based on the principle that torque developed 
at the motor shaft is proportional to both the rotor current 
and the square of the terminal voltage. Therefore, if the 
terminal voltage is reduced to a given level to correspond 
with a desired torque value, then motor current is limited. 
Thyristors are used to control power turn-on to the ac 
motor, and voltage is reduced by not firing the thyristors 
until some angle past the source voltage zero. This effec- 
tively reduces the average voltage across the motor termi- 
nals to a value less than the source voltage. Firing control 
is by one of two means: open loop or closed loop (2, 5, 31). 

Without feedback, the open-loop systems are the sim- 
pler designs, and the terms voltage ramp and current ramp 
are often used to describe these systems. The controller 
element (G in figure 14.9) brings the voltage or current up 
to maximum in a predetermined time period independent 
of the motor conditions. The maximum value of current is 
not limited by the starter, but rather by the line voltage 
and motor characteristics. This limits the use of open-loop 
control to motors or power systems where the starting 
current does not create problems. 

In the closed-loop systems, some reference signal or 
feedback from the motor (H in figure 14.9) is compared 
with a reference to adjust the controller output. The result 



353 



is more control over the motor starting characteristics and 
allowances for loading or operating conditions. In the case 
of static belt-conveyor starters, and feedback signal typi- 
cally corresponds to either motor line current or motor 
speed. The line-current feedback control schemes can be 
divided into current-limit or current-regulated types. 
Motor-speed feedback techniques are also called linear 
acceleration or tachometer control. 

The basic current-limit scheme is shown in figure 

14.18. Feedback control is used to compare an adjustable 
reference signal (the current-limit setting) with a signal 
that represents motor current. Motor current is monitored 
by current transformers, preferably in all three lines (2). 
The torque being developed at the motor shaft is then 
computed by the starter circuitry from the current being 
used. As a result, the current is limited to a preset value 
by the thyristors, and the resulting belt acceleration is 
nonlinear (almost logarithmic) and is a function of belt 
load. When the belt is empty, a full-voltage start can occur. 

In current-regulated control, the problem of a full- 
voltage start is removed by the addition of a second 
reference supplied to the control-system summing point. 
This is an adjustable ramped reference signal, which is 
usually a linearly increasing voltage with time. The 
starter circuitry now restricts motor current to rise over a 
preselected time period (the acceleration-time setting in 
figure 14.18) to the current-limit setting. Motor accelera- 
tion is smoother than current-limit control but is still 
somewhat nonlinear. 

In basic linear-acceleration starters, a tachometer 
generator is placed on the motor shaft, and its output 
provides either a feedback signal proportional to motor 
speed, which is compared with a ramped reference, or a 
feedback signal proportional to motor acceleration, which 
is compared with constant preset reference. As a result, 
motor voltage is adjusted to limit current, such that a 
specified rate of acceleration is obtained. Belt acceleration 
is linear and rather constant regardless of load, this is, 
whether the belt is empty or full. Some models combine 
the linear-acceleration feature with overriding current- 
limit and current-regulation control, as shown in figure 

14.19. Thus, the starter tries to linearly accelerate the 
motor, but the current-regulation control keeps the rate of 
current rise within preset limits and the current-limit 
control keeps the maximum value of current below a 
certain level. Controlled deceleration is also available in 
some starters, but the deceleration time must be longer 
than that for the drive mechanical inertia. 

The start-stop circuitry is often relay controlled. (This 
is not shown in figures 14.18 and 14.19.) When off, relay 
contacts clamp the thyristor gates and the input of the 
control-system amplifier. Under a start command, the 
relay contacts sequentially unclamp the gates, allowing 
the control system to start the acceleration cycle (2). 
Because the thyristors do not physically disconnect the 
load, a circuit breaker is also provided on the incoming 
line for this purpose as well as for short-circuit protection. 

Control System Design Considerations 

When starting an induction motor on a belt drive, 
three major areas need to be considered: the power system, 
the motor, and the mechanical equipment (speed reducers, 
the belt, and so forth). From the standpoint of power- 
system protection and optimum operation, the motor 
should be started with as little current as practical to 



Current- 
limit 
setting 




(^ — (§H>- 



Firing 
control 



Acceleration- N 

time ' 

setting v 



! Ramp j 
"generator' 



TTT 

L-i L 2 L 3 



Figure 14.18.— Simplified diagram of current-regulated 
static belt starter. 



Motor 



Current- 
limit 
setting 

Acceleration- 




,, Amplifier 



time ?\ » 
setting v_^ 



To other 
thyristors 

Jill 



coupling 

i 
Tachometer Q 



irmg 
control 



HT 



Tachometer 

signal 
processing 



L, L-, 



Figure 14.19.— Simplified diagram of linear-acceleration 
static belt starter. 



minimize the overcurrent and undervoltage effects on the 
power system. With motor protection and operation taken 
as the priority, the philosophy would be to bring the motor 
up to full speed as quickly as possible. The reason here is 
to prevent insulation breakdown caused by rotor overheat- 
ing. The optimum situation for protection of the mechan- 
ical equipment would be a smooth, easy acceleration of the 
belt. This reduces the wear and tear on the gears and the 
belt from excessive or uneven torque. 

Unfortunately, because the design considerations are 
mutually exclusive, all three areas cannot be completely 
protected. Protection of the power system and mechanical 
components calls for an extended low-level starting cur- 
rent, whereas protecting the motor from overheating re- 
quires a rapid-rise high-level starting current. The area 
that is given priority helps determine which control sys- 
tem should be used (31). 

When solid-state belt starters were first introduced, 
protection of the power system and conveyor components 
was the primary concern. Thus, early starters used a 
current-regulated or linear-acceleration scheme for 
smooth starting of the belt and an overriding current limit 



354 



to protect both the power system from overcurrent situa- 
tions and the conveyor components against overtorque 
problems. Either control system gave the same basic 
static-starter advantages given previously. Linear- 
acceleration controls had a special advantage in installa- 
tions where belt length was constantly changed, such as 
for panel belts, as little or no adjustment of the starter 
circuitry was necessary. With current-regulator control 
only, adjustment may be required for any conveyor change, 
but current-regulator controls do have the advantage of 
simplicity. Most manufacturers offered both options in 
their equipment. 

Many of these early static starters worked well, but 
some problems occurred because of initial design flaws. 
Foremost was the situation where either the current limit 
was set too low or the belt was overloaded. In either case, 
the motor would not reach full speed before the rotor 
overheated because of the somewhat limited acceleration 
imposed by the current limit, and either the thermal 
protection of the motor or a maximum current time limit 
would shut down the motor. Another common problem 
area was strictly associated with linear-acceleration mod- 
els: high maintenance requirements of the tachometer 
generator and its wiring. 

These problems can directly produce conveyor belt 
downtime and, in turn, a large loss-of-production cost for 
the mine. Thus, a recent trend with many manufacturers 
has been toward more reliability and a simpler control 
system, and the common design is an open-loop voltage 
ramp scheme without a current limit. Typically, the volt- 
age starts at roughly 60% of the line voltage then gradu- 
ally increases to 100% in a preset time limit. Because 
there is no current-limiting capability, thermal overload 
sensing in the motor is almost a necessity, and simple 
circuitry for tripping the main circuit breaker when the 
overcurrent situation has exceeded a specified time limit is 
usually recommended. This design is reported to have a 
lower breakdown rate than other more complicated sys- 
tems, and it also allows for full locked-rotor torque and full 
locked-rotor current when needed by the conveyor. How- 
ever, the simplicity does negate some of the advantages of 
soft starting, particularly limiting starting currents and 
voltage drops. 

Motor Designs 

There are four concerns when selecting a suitable 
motor and motor characteristics: providing sufficient 
starting torque, drawing symmetrical line current with 
minimum disturbance to other equipment, optimizing the 
thyristor characteristics, and providing an acceptable 
stress level on motor insulation under any condition (39). 
In terms of characteristics, NEMA design B appears to be 
a good match for linear-acceleration control because of its 
greater locked-rotor time, and NEMA C seems best suited 
for open-loop and current-regulated applications from its 
greater starting torque at a given current (31). However, 
the main consideration is thermal because the motor must 
accelerate to full speed for normal operation. Energy 
continues to be stored in the motor until the internal fan 
velocity can dissipate the heat faster than it is generated. 
Hence, NEMA C motors are often recommended for all 
control schemes because less energy output is needed to 
provide the same output starting torque. Some manufac- 
turers have even recommended the NEMA design D be- 
cause they feel the advantage of its very high starting 



torque outweighs the disadvantage of its inefficient full- 
speed characteristics. 

Some motor manufacturers are producing squirrel- 
cage machines specifically designed for solid-state start- 
ing. They have higher quality insulation, larger fans, 
larger rotors, and are built on a larger frame per given 
horsepower. These motors can have an energy-in-heat-out 
balance at 75% of full speed, whereas for the standard 
motor, it can be as high as 90% of full speed (39). 

Experience has shown that conventional thermal- 
overload relays do not provide adequate protection during 
the long starting times caused by static control (2, 30). 
Overtemperature detection devices installed in the motor 
are the best way to provide thermal protection. Of these, 
solid-state temperature sensors, installed internally in the 
motor windings, with a load-current detection backup, 
appear to provide the best technique (2). 

Thyristor Configuration 

Both thyristor configurations, unidirectional and bi- 
directional, are available for static belt conveyor starters, 
but there are difficulties with unidirectional control. Some 
problems are caused by the higher harmonic content that 
unidirectional control adds to line current when in the 
control mode. When applied to static induction motor 
starters, the harmonics tend to create excess heating but 
mainly during acceleration. If sufficient cooling time is 
allowed between starts and acceleration times are mini- 
mized, the heating problem is reduced considerably but 
not eliminated (5). 

However, the main difficulty with unidirectional con- 
trol is allowing dc to flow in the motor, and two situations 
illustrate the problem (5). During a motor ground fault, 
the diodes can rectify current, and dc can flow through the 
grounding resistor and grounding conductor, and not be 
detected by the zero-sequence ground-fault relaying. Thy- 
ristor failures are rare but are always in the shorted mode. 
If thyristor fusing is not available, there also exists a 
low-impedance path for dc through the motor windings 
and the other two diodes. Either case is a problem whether 
the starter is on or off. 

With the harmonic and dc difficulties presented by 
unidirectional control, it appears that bidirectional control 
is better despite its high cost. Some additional advantages 
are gained (39). Two thyristors are always in series from 
line to line, and hence, less stress is given per thyristor by 
transient and long-term overvoltages. Because of symme- 
try, smoother control is provided under varying conditions, 
and finally, motor windings and cable power conductors 
are near neutral potential when the thyristors are off. 

Firing Circuits 

When discussing the basics of thyristors, a single gate 
pulse was shown to fire the device (figs. 14.3, 14.20A). 
However, in practice, there are many instances where a 
single gate pulse would fail to fire the thyristor, for 
example, if there is low ambient temperature or low line 
voltage or when the devices themselves are old (39). Two 
types of firing systems overcome this problem: sustained 
pulse (fig. 14.20B) and multi-pulse (fig. 14.200. Generally, 
a pulse transformer is used in the multipulse system as 
isolation between the power circuit and the control circuit. 
Numerous pulses, each capable of turning the thyristor 
on, are applied to the gate to keep the thyristor on during 



355 



'g' 


I 















i 


. 


i 







oc 


IT 


27T OC + 2TT 


3ir 





Ir.A 



cut 



_L 



■cut 





'g* 



TT 



2TT QC+2TT 




3TT 




cut 



TT 27T CC+2TT 

Figure 14.20.— Types of thyristor firing pulses. 



3TT 



its intended conduction period. The other technique, also 
termed the dc firing system, maintains a continuous gate 
current during the desired conduction period. This helps 
ensure that the thyristor will turn on, and turn on 
completely by continuous stimulation of the gate. 

Two parameters of the firing pulse are important: the 
level of the current and the current duration. If the pulse 
is only of minimal current and duration, it may cause 
conduction only in a limited area of the thyristor or no 
conduction at all. In particular, a small conduction area 
coupled with a high rate of change (di/dt) of anode- 
to-cathode current can result in concentrated heating and 
possible failure of the thyristor. To ensure conduction of 
the whole interface of the thyristor, a technique called 
hard firing is used in conjunction with either dc or 
multipulse firing. It consists of a high-level initial pulse 
with a steep wave front and enough duration to operate 
the thyristor at near the maximum input power level of 
the gate. The high initial gate current floods the conduc- 
tion region of the thyristor and turns it on completely. 
After the device is conducting, a sustained dc pulse or 
multipulse keeps the thyristor on even if the anode- 
to-cathode current approaches zero. This prevents thyris- 
tor turnoff during the critical end of the conduction cycle. 
Hard firing, thus, provides more consistent operation and 
reduced failures of thyristors, and allows the use of off- 
the-shelf devices (9, 39). 

Thyristor Ratings and Protection 

Thyristors are susceptible to damage from overvolt- 
ages, overcurrents, and rapid changes in voltage (dv/dt) or 
current (di/dt) with respect to time. Because these are 
rather common occurrences in mine power systems, thy- 
ristors should be protected from each of these phenomena 
if premature failure is to be prevented. The best way to 
protect the thyristors against the effects of high di/dt, 
which has already been mentioned, is proper design of the 
firing pulse. Two avenues are typically used to provide the 
other protection, thyristor ratings and protective circuitry, 
and a typical protection arrangement is shown in figure 
14.21. 



I 2 t 

fuse 



0.25AF 



50 A 



MOV 



^ 



MOV 



/ Optional 
Q indicator 
/T^ lamp 



- t f 



Gate pulse 
to control 
(di/dt) 



Figure 14.21.— Thyristor protection for static belt starters. 



It is standard practice to select a thyristor voltage 
rating at 2.5 times the nominal line-to-line system voltage. 
For bidirectional control, this is quite adequate consider- 
ing the 5-pu utilization transients discussed in chapter 11. 
However, the thyristors themselves are a source of abnor- 
mal transient overvoltages, and some engineers prefer to 
use a device voltage rating no less than 3 or 3.5 times the 
system voltage (5, 39). 

Calculation of the required thyristor continuous- 
current rating from the load current should be easy if the 
thermal properties of the heat sink are known. Yet during 
starting, which is the worst stress case, the device is called 
upon to deliver much more current. The conventional 
approach for overload protection of the thyristor is a 30-s 
(overload) rating of 300% continuous current. Some engi- 
neers do not believe this is adequate, because the induc- 
tion motor starting current may be as high as six to eight 
times the running current; hence, certain manufacturers 
have selected thyristors with 300% of the continuous- 
current requirements and a 30-s overload rating at 500%. 
Others feel that it is better to have an overload rating 



356 



based on horsepower, 25 s at 600%, or 100 s at 300% of 
continuous current (39). Thus, the thyristor current-rating 
selection is not a straightforward matter, and special 
consideration is sometimes required for individual cases 
(5). 

Some short-circuit protection is afforded by oversizing 
(300%) the continuous-current rating. However, past prac- 
tice was to provide additional protection with semiconduc- 
tor protection or I 2 t (let-through energy) fuses with one 
fuse in series with each device. The philosophy was that, 
even though these fuses were expensive, the replacement 
cost for the thyristor was even more so. The fuse 
continuous-current rating was matched to that of the 
thyristor, but the I 2 t rating was less than the thyristor I 2 t 
capability. There were instances where the thyristor 
continuous-current rating was set at 150% of needed and 
then the I 2 t fuse was matched (39). However, this proxim- 
ity was found to lead to nuisance fuse activations, espe- 
cially during acceleration. 

Recent practice for overcurrent protection of the thy- 
ristors is to eliminate or oversize the I 2 t fuses. When the 
fuses are eliminated, the thyristors are sufficiently over- 
sized to withstand most long-duration situations, and 
thermal-overload sensors are used on the thyristor heat 
sinks. In cases where the fuses are used, the fuse current 
limit is set very near the failure point of the thyristor. 
Both of these schemes are intended to reduce starter 
downtime due to blown fuses, even to the point of sacrific- 
ing the main power thyristors. The general feeling among 
mine operators is that the production saved by eliminating 
false fuse tripping is worth an increased liability of the 
thyristors' failing. This situation has become justified 
because of increased production cost associated with belt 
downtime and the decreased cost of the thyristors. 

A high rate of forward-voltage increase can turn a 
thyristor on, even with a zero gate current (9). Inductive 
loads, such as induction-motor belt drives, can present 
such dv/dt problems (5). The common solution is RC 
snubber networks across each pair of devices as in figure 
14.21. Typical values for R s and C s are 50 fi and 0.25 fiF, 
but values must be selected so that the dv/dt allowed is 
less than the minimum specified by the device (around 100 
V//ts). To ensure that transient overvoltage does not de- 
stroy a thyristor, metal oxide varistors (MOV's) across each 
device are sometimes specified. The maximum crest al- 
lowed by the MOV is coordinated with the maximum 
voltage rating of the thyristor, allowing the usual 20% 
safety margin. 

The preceding paragraphs have not only introduced 
the basics of solid-state motor control but also indicated 
the special difficulties of applying static-starter concepts 
to mine belt conveyors. This is one firm example of the 
benefit of technology to industry. Even with the increased 
internal complexity, the end result has been an overall 
increase of belt-conveyor system reliability. 

STATIC PROTECTIVE RELAYING 

Most engineers define the term relay as an electrically 
controlled, usually two-state, device that opens and closes 
electrical contacts to effect the operation of other devices 
in the same or another electrical circuit (15). Historically, 
one important relay use has been the protection of people 
and electric circuits from electrical hazards. The operation 
of an electromagnetic protective relay was presented in 
chapter 9 but is repeated in figure 14.22 (36). This partic- 



Station bus 



CT 



u - 

b 
c — < 


1 1 




1 


i — 




( 




1 










nr 


JL 

nr 

nda 








4_- 

i 
i 
i 
i 

i 
i 
i 


PT 


Circuit 
breaker 


C Trip 
C coil 










A ° 








[_ 


-o o- 

\ i 










£■ 








i 

eco 














s 

p 




Protected 
circuit 




Relay 


r y 












otential 
bus 



Figure 14.22.— Protective-relay connections. 



ular relay uses two actuating quantities (voltage and 
current) that directly affect the status of the relay con- 
tacts. Whenever current and/or voltage exceeds a prede- 
termined level, the current sensor and/or voltage sensor 
(that is, CT and/or PT) outputs cause the relay to close its 
contacts through electromagnetic attraction or electro- 
magnetic induction. The closed contacts permit current to 
flow through the trip coil, tripping the circuit breaker. 
Electromagnetic-attraction relays in common use are the 
solenoid, clapper, and polar types (38). The typical 
electromagnetic-induction relay is the induction disk. 

The basic concept in the design of solid-state relays 
(again also called static relays) is to replace the mechani- 
cal contact device with a solid-state device. The solid-state 
device is inserted in the trip coil circuit and controlled by 
the sensor circuit. When unactuated, the solid-state device 
acts as a very large resistance in the trip coil circuit, 
limiting the current through the trip coil to a very small 
value (known as leakage current), which is incapable of 
tripping the associated circuit breaker. When actuated, 
the solid-state device acts as a very small resistance, 
which allows ample trip coil current, thus tripping the 
circuit breaker. The solid-state devices commonly used are 
the transistor, the thyristor, and the triac. 



OPERATION OF SIMPLIFIED SOLID-STATE AND 
HYBRID RELAYS 

The electromagnetic relay of figure 14.22 is repre- 
sented schematically by figure 14.23, where the current 
and voltage input have been replaced with a manual push 
button, and the trip coil (load) circuit is supplied with a 
120- V, 60-Hz power source. The static relay differs from 
this in that the contents of the dashed box of figure 14.23 
are replaced by the semiconductor device shown schemat- 
ically in figure 14.24 (16). The load current now flows 
through the common terminal of the semiconductor de- 
vice, which is common to the trip (contact) and sensor 
(control) circuits. 



357 




Figure 14.23.— Simple electromechanical relay. 



Semi- 
conductor 
contact 
device 

\ 



Control 

current 

input 



Load 




Control 



Common 




Control 
current 



Output 
( contact ) 



Figure 14.24.— Simple static relay. 



If the static device is an npn transistor, the circuit of 
figure 14.25A results. Since the transistor is a current- 
controlled device, if zero input (control) voltage is applied 
to the transistor base, no current will flow into the base. 
Because base current is required for current flow from 
collector to emitter, no current will pass through the load. 
Hence, the contact-circuit voltage supply will not be 
dropped across the load, and the voltage must therefore be 
across the collector-emitter terminals of the transistor. A 
positive input voltage applied to the transistor base causes 
a positive current to flow into the base. This base current 
is the controlling factor that permits a large collector- 
to-emitter current to flow through the transistor. When 
sufficient base control current is supplied, the collector 
current will increase until essentially all contact-circuit 
voltage is across the trip coil (load). Thus, the switching 
characteristics of figure 14.25B are obtained. 

A drawback to the circuit in figure 14.25A is the lack 
of electrical isolation between the control and the trip 
circuits. One isolation method is to use the photon-coupler 
circuit shown in figure 14.26 (16). Here, the transistor 
performs the same function as that in figure 14.25A except 
that it is light controlled, with a high-intensity light (from 
the LED) acting as a large base current. The light im- 
pinges on the phototransistor base region, allowing cur- 
rent to flow from collector to emitter. Without an input to 
the LED, no light is produced and consequently no 
collector-to-emitter current flows. 

Replacing the transistor with a thyristor results in 
the circuit of figure 14.27 (16). This device has been 
described earlier in the chapter; thus, its operation here 
should be clear. 

The triac, which is a contraction for triode ac semi- 
conductor, is a bidirectional solid-state device that acts 



'load 



'input 



Current- 




- Load 1 


limiting 
resistor JC 

t Control V 

1 1 


Contact I 
circuit 
voltage 
(V) « 

ransistor 

1 



'load 



'input 



A Circuit 



Control Control 

voltage voltage 

applied removed 

B Switching waveforms 



Figure 14.25.— Transistor used as relay. 



Current- 
limiting 
resistor 
I — VW\r— 




Figure 14.26.— Optical transistor as relay. 



V 



load 



^WvV^ 



v input| Control 



Load 



V, 



scr 



rms 



Figure 14.27.— Thyristor used as relay. 



like two thyristors connected back-to-back as shown in 
figure 14.4A. The triac provides full-wave voltage control 
in one solid-state structure with only one gate control, as 
shown in figure 14.28A. The load-voltage characteristics, 
as shown in figure 14.28B, are very much like those 
exhibited by the thyristor. The single structure has heat- 
dissipation limitations, restricting the triac to small- 
current applications (9). 



358 



V. 



i i 



v 



load 



KAAAA 



Triac 



MT2 



^MArTlM 



' input 



MT1 



Load 



load 



'triac 




V ( 



triac 



<M 



KEY 
MT1 Main terminal 1 (common) 
MT2 Main terminal 2 ( load) 
G Gate (control) 



V; 



input 



1 I 
I i 

Triac self- 
commutates 



!/ 



¥\ 






1 

1 


Control 
voltage 
applied 


Control 
voltage 
removed 



Figure 14.28.— Triac used as relay. 



The output and input circuits of the triac and thyris- 
tor relays again have a common terminal between them. 
Where input-output isolation is required and a solid-state 
relay is also desired, hybrid relays are commonly used. 
Hybrid relays employ a relay with a low-power operating 
coil in either the input or output stage and a solid-state 
device functioning in the other stage. The hybrid relay 
shown in figure 14.29A has the reed relay in the input 
stage (8). Switching the reed relay activates the gate- 
control input, which fires the triac. The hybrid circuit of 
figure 14.29B uses a solid-stage input stage that can react 
to the power network conditions. When a predetermined 
threshold is reached, the solid-state stage sends a current 
through the relay coil that will pick up the relay contacts. 

Reed relays alone, when compared with static devices, 
have the advantages of simplicity, low cost, high reliabil- 
ity, and low maintenance. The best reed contacts are 
mercury wetted and provide bounce-free operation. Dry 
reed relays with heavy-duty silver contacts will carry 2 kW 
with a maximum of 30 A. They can withstand 5 G (50 m/s 2 ) 
when mounted correctly. The characteristics of electro- 
magnetic (attraction and induction types), reed, transistor, 



and thyristor (including triac) relays are summarized in 
table 14.1 to allow a quick comparison (37). This table 
shows that static relays outperform electromagnetic relays 
in almost every case. 

Before a detailed comparison of static and electrome- 
chanical relays can be made, it is necessary to determine 
whether static relays are capable of performing the same 
functions as electromechanical relays. The latter are usu- 
ally classified according to the actuating quantity (see 
chapter 9), and the most common types used in mining are 
overcurrent, overvoltage, and undervoltage. The overcur- 
rent or overvoltage relay is designed to trip when a 
predetermined current or voltage threshold value has 
been exceeded. The undervoltage relay is designed to trip 
whenever the measured voltage falls short of a predeter- 
mined level. Static relays can be designed to respond to 
these actuating quantities as well as many others (37). 

As an example of static operation, consider the over- 
current relay shown in figure 14.30. The ac from the CT is 
forced through a resistor, which produces an alternating 
voltage across the input rectifier bridge. This voltage is 
rectified by the bridge and applied to the series RC 



Table 14.1.— Typical electromechanical and static relay characteristics 



Parameter 



Electromechanical 



Static 



Electromagnetic 


Reed 


Transistor 


Thyristor 


1-3 


0.1-3 


0.010 


0.040 


30 


<20 


50 


1,200 


10-30 


7-200 


5,000 


30,000 


5 


1.0 


1 


5 


10 


1-2 


0.020 


0.050 


<6 


<6 


1 


1 


10 7 


10 7 


( 1 ) 


( 1 ) 


-5-70 


-5-55 


-20-100 


-20-100 


Yes 


Little 


No 


No 


Yes 


No 


No 


No 



Input W.. 

Switching capacity W.. 

Power gain 

Continuous carrying current A.. 

Delay ms.. 

Number of contacts 

Maximum operations 

Ambient temperature range °C. 

Affected by vibration 

Affected by corrosive atmosphere 

1 No limit. 



359 



i — WW — f 



r\ 



I 



Reed 
relay 



Triac 



& 



6 

Load 

o 




A Solid-state output B Solid-state input 

Figure 14.29.— Hybrid static relays. 




Relay 
supply 
voltage 



Battery used as 
pickup reference 

Figure 14.30.— Simple overcurrent static relay. 



elements in the base-emitter circuit. When the capacitor 
voltage exceeds the battery voltage in the base-emitter 
circuit by a sufficient amount, the transistor begins to 
conduct current. Thus current picks up the trip coil, 
tripping the circuit breaker. A time delay is achieved in 
this circuit by virtue of the series RC time constant 
associated with the base-emitter circuit. The other actu- 
ating quantity operations can be described in a similar 
manner. 



STATIC AND ELECTROMECHANICAL RELAY 
COMPARISON 

A large amount of literature has been devoted to the 
advantages and disadvantages of static relays. The follow- 
ing detailed comparison summarizes this literature and 
discusses the advantages and disadvantages associated 
with the differences between static and electromechanical 
operation. A listing of the references used follows each 
topic of the discussion. 

Relay Speed 

Table 14.1 shows that static relays are much faster 
than electromechanical relays. This speed difference is 
because there is no mechanical motion in solid-state 
devices. High speed is advantageous since it gives the 
static relay the capability to turn on at zero-voltage 
crossing. This eliminates much of the radio frequency 
interference (RFI) caused by capacitance switching or 
prestrike that is found with mechanical opening and 
closing of circuits with large capacitance values. Control- 
ling RFI is important because digital circuits are easily 
affected by such interference. RFI is not present when 



static devices open, since for example, the thyristor and 
triac always turn off at zero current. The variability in the 
closing and opening time precludes this type of control in 
electromechanical relays. Reset times are also much 
shorter for static relays. An extremely inverse electrome- 
chanical relay may require 15 to 25 s to reset, while a 
static relay resets in a few cycles (1, 8, 18, 22, 26, 37). 

Relay Power Requirements 

As shown in table 14.1, static relays require far less 
input power than electromechanical relays. Input power 
delivered to any one static device may come from a wide 
range of voltages, including ac and dc. Electromechanical 
relays are normally limited in the voltage range they can 
accept for proper operation and require different coils for 
ac and dc operation. Static relays are also capable of using 
a range of ac and dc power supplies in their tripping circuit 
(3, 18). 

Heat production in the solid-state relays results from 
the 1.0- to 1.5-V drop across the device in the on position; 
thus, power consumption is limited by the current re- 
quired for tripping operation. Static relays used for 
protective-relaying applications usually require only 5 A, 
which is insufficient to produce enough heat for concern. 
Relays carrying between 5 and 25 A, usually referred to as 
power relays, may require an auxiliary heat sink or special 
mounting arrangements. Above 25 A, a relay becomes a 
contactor, even though its function does not change; hence, 
a static relay above 25 A is considered a solid-state 
contactor. Electromechanical relays require no heat sink 
since, in an underrated condition, their contact resistance 
measures in the milliohm range (1, 8, 18). 

Temperature 

Typically, the maximum allowable continuous junc- 
tion temperature for thyristors and triacs is 120° C, while 
electromechanical relays may be used in ambient temper- 
atures that exceed this value. Temperature has little effect 
on the electromechanical characteristics; however, solid- 
state devices exhibit different characteristics at elevated 
temperatures. If subjected to these temperatures for long 
periods, their characteristics may even change perma- 
nently. Short-range thermal compensation can be accom- 
plished by the use of thermistors in the design or, more 
commonly, differential-stage construction. The long-range 
effects of elevated temperature are guarded against by 
preassembly and postassembly heat soaking and testing 
(1, 18, 37). 



360 



Power Transient Response 

The term power transients refers to transients on the 
power system being monitored and transients on the trip 
circuit control power. Static relays are fast-responding 
devices, and they rarely trip on power-system transients 
unless set to do so. Electromechanical relays, however, may 
trip erroneously owing to either ratchetting or overtravel. 
As was discussed in chapter 10, mechanical inertia causes 
overtravel of the induction-disk or induction-cup relay 
even after the actuating quantity has been removed. If the 
relay is not set properly, overtravel may cause the unde- 
sired tripping of a backup relay, thus isolating a sound 
part of the power system being protected. Static relays 
have negligible overtravel, which precludes this type of 
error and makes it possible to have closer time settings on 
backup relays to provide faster protection. Table 14.2 
shows a comparison of the time-margin calculations for a 
backup time-overcurrent relay. These calculations assume 
that the circuit breaker operates within 3 to 8 cycles after 
pickup has been exceeded in the primary relay. A toler- 
ance margin is allowed for the normal variation in the 
time-current curves. The smaller tolerance margin used in 
the minimum time calculation corresponds to establishing 
the trip time by testing, while the larger margin assumes 
a time selected using the time dial settings on the relay. 
For practical purposes, a 0.3-s margin is used for the 
electromechanical and a 0.2-s setting for the static relay. 
The difference is due to the negligible overtravel of the 
static device (19, 22). Ratchetting is the accumulation of 
overtravel due to successive faults or sequential starting of 
motors. Static relays do not exhibit ratchetting because of 
their negligible overtravel and short resetting time (22). 

Table 14.2. — Time-margin comparison between 
electromechanical and static relays, seconds 



Operation 


Electromechanical 

Min 1 Max 2 




Static 


Min 1 


Max 2 


Breaker time 

Overtravel 

Tolerance margin 


0.05 
.10 
.07 


<0.13 
<.10 
<.17 


0.05 
.01 
.07 


<0.13 
<.01 
<.17 


Total 


.22 


£.40 


.13 


<.31 



Trip circuit control-power transients may cause false 
operation of an electromechanical relay if that device 
relies on the control power to keep it closed under normal 
circuit conditions. Voltage spikes on the control-power 
circuit can also cause unwanted tripping of static relays 
from two separate situations. If the transient is of suffi- 
cient magnitude, the solid-state device may break down, 
permitting conduction. This problem has to a large extent 
been overcome by triacs with blocking voltages of 1,000 V 
and higher. However, if the voltage spike has a high rate of 
rise, the inherent capacitance of the thyristor or triac will 
allow a momentary current to flow (that is, i = Cdv/dt). 
With sufficient magnitude, this current will act as a gate 
current and turn on the device. This dv/dt problem is 
similar to that discussed for static belt conveyor starters 
and is being overcome by including filters capable of 
absorbing the energy contained in these transients. On 
the other hand, dv/dt current is not a problem with 
electromechanical relays. Breakdown voltages of these 
relays are normally many times their operating voltage 
because of the air gap between the mechanical contacts (1, 
18, 26, 30). 



Mechanical Nature 

Electromechanical units have contacts that close me- 
chanically and electrically, while static units close only 
electrically. The differences that result from the construc- 
tion of the two types of relays can be described by compar- 
ing the relays when the contacts are open or closed, and 
when the contacts are in the act of opening and closing. 

Unlike static relays, which are usually encapsulated 
in epoxy, open electromechanical contacts are exposed to 
the surrounding atmospheric conditions. Settling dust can 
act as an abrasive agent and cause premature contact 
failure, and chemical contaminants promote oxide-layer 
production, which can also increase contact resistance. 
However, the physical separation of the open contacts 
provides a very large resistance in the trip coil circuit, 
which cannot be matched by a solid-state device. The 
leakage current that flows through the solid-state device 
when in a high-impedance state is on the order of 10 to 20 
mA. Static relays should not be used where this current 
magnitude cannot be tolerated (16, 18, 26, 29). 

When electromechanical contacts are closed, they 
normally have only a few milliohms of resistance to 
control-power current, while thyristors or triacs provide a 
1.0- to 1.5-V voltage drop. If the contacts become so 
degraded that contact resistance causes excessive heat 
generation, the contacts may weld closed. Moderate over- 
loads can be tolerated for short periods. On the other hand, 
when solid-state relays are underrated for a particular 
application, they will fail catastrophically. Where more 
than one contact closure is necessary, additional contacts 
are inexpensive and easy to add to electromechanical 
relays, but they can only be added to static units by a 
costly duplication of solid-state circuitry portions. Simi- 
larly, interlocks and remote indicators are also more 
expensive in static relays; in fact, these tasks may be 
performed more economically by using digital logic on the 
control side (8, 18, 26, 29). 

The input-output coupling of electromechanical relays 
provides a very large isolation resistance between input 
and output. Hybrid relays can match this resistance, while 
straight static devices can approach a value of between 
10 10 to 10 11 if the photon coupler isolator shown in 
figure 14.26 is used. 

Although a contact is in the act of closing or opening 
for only a small portion of its life, this movement is the 
essence of the difference between static and electrome- 
chanical relays. It adds the problem of potential arcing to 
the operation of electromechanical relays: arcing does not 
occur in static relays because of the nature of the solid- 
state contact. High-inductive circuits will usually cause 
contact arcing during separation, and arcing can pit the 
contact surfaces, giving a higher value of contact resis- 
tance. Arcing is particularly dangerous in explosive envi- 
ronments. Closing contacts are subject to contact bounce, 
which can also cause arcing (8, 16, 18, 26, 29). 

Testing electromechanical relays by listening for con- 
tact closure is a simple method of inspection. Early static 
units had no equivalent test procedure, but more modern 
static relays have pushbutton testing that uses a target 
(flag) to signal proper operation (8, 26, 34). 

Both moving and stationary mechanical contacts are 
subject to nuisance tripping due to external vibrations, 
and this is especially bothersome with the portable power 
equipment used in mining. The mechanical nature and 
diverse mechanical configurations of electromechanical 
relays make their response difficult and costly to predict. 



361 



The epoxy resin covering, physical structure, and struc- 
tural arrangement of static relays make them much less 
susceptible to seismic disturbances (15, 29). 

Versatility 

If there is any one characteristic, other than reliabil- 
ity, that ensures an ever-increasing role for static relays, it 
is their versatility. They can duplicate the time-current 
characteristics of electromechanical relays described in 
chapter 9 and also provide functions that heretofore were 
impossible. For instance, static directional-overcurrent 
relays are feasible and provide unique characteristics for 
improved operation (11); new static negative-sequence re- 
lays are capable of tripping at only 10% negative-sequence 
component content, where electromechanical relays are 
limited to tripping levels of 18% negative-sequence com- 
ponent (22, 32). Many static relays are compatible with 
digital logic (8), and it is expected that the capability and 
versatility of this mode will push static relaying in direc- 
tions not even considered previously. More will be said 
about this when the future of static relaying is discussed. 

Current Transformer Burden 

CT burden has been defined as the external load 
applied to the secondary of a CT. As noted in chapter 10, 
high CT burdens can cause core saturation because of the 
magnitude of the voltage developed at the CT secondary 
terminals. Saturation causes a burden reduction since, 
under deep saturation, the burden approaches its dc resis- 
tive value. This burden change modifies the relay charac- 
teristics and can cause incorrect operation (36). 

The possibility of CT saturation is of particular im- 
portance when low-ratio CT's cause large secondary cur- 
rents (large secondary voltages) and when large burdens 
from additional secondary loads cause large secondary 
voltages from even moderate secondary currents. Hence, 
relays presenting lower burdens are clearly advantageous. 
Table 14.3 compares the burdens presented by a typical 
electromagnetic-induction overcurrent relay with its static 
counterpart when set on the minimum tap setting. The 
static relay burden includes that of the relay current- 
measuring portion and the power-supply portion because 
the relay receives its operating power from the CT. The 
burden of the current-measuring portion of the static relay 
is actually much less than that for the electromechanical, 
but note that the burden is greater than the electrome- 
chanical burden at pickup. The lower static burden has the 
advantages of allowing: 

• More relays to be connected in series, 

• The lowest tap of a multiratio CT to be used, 



• A step-up auxiliary CT to be used for greater 
sensitivity, and 

• More auxiliary burdens to be connected without CT 
saturation. 

Another advantage is that the low CT burden will permit 
the use, without saturation, of a less expensive CT that 
has no relay classification (19, 34). 

Accuracy 

In general, static relays are more precise than their 
counterparts for reasons other than those already men- 
tioned. Common induction-disk relays are adjustable 
through 11 discrete tap settings in one range only. A 
typical static time-overcurrent relay can have an additive 
tap-block arrangement, which permits numerous tap set- 
tings in more than one range. At low multiples of pickup, 
typically below 1.5 times pickup, induction-disk relays 
produce very low torque levels, which can make time 
settings unpredictable. Static relays are reported to work 
reliably down to 1.1 multiples of pickup. The frequency 
content of the signal applied to a relay is another source of 
error. Static relay response in the band of interest is 
uniform, while induction-disk relays show a distinct vari- 
ation of response with frequency. The uniformity of static 
relay frequency response should make the performance of 
static relays more predictable (19, 22). 



STATIC RELAY MINING APPLICATIONS 

An outstanding application of solid-state relaying to 
mining has been in ground-check monitoring, especially 
when connected to portable or mobile face equipment. 
Because the success of these devices has been adequately 
covered in chapter 9, nothing more will be said here. Static 
protection relays have been applied successfully in U.S. 
mines for the following common ac distribution areas (7, 
10): 

• Line short circuit and overload (device 50/51), 

• Undervoltage (device 27), 

• Phase loss and phase sequence (device 47), 
Zero-sequence ground fault (devices 50G and 51G), 



and 

• Grounding-resistor overvoltage (device 59G). 

In all these instances, the replacement of electromechan- 
ical devices with static devices has led to greater environ- 
mental resistance, easier adjustment and testing, more 
sensitive ground-fault protection, and lower maintenance 
time (10). 



Table 14.3.— Comparison of induction-disk and static time-overcurrent relay burdens to 

impedance in ohms 


a current transformer, magnitude of 




Relay range, 1 0.5 to 4 A 






Relay range, 1 1.5 to 12 A 




Relay 

1X 


3X 10X 


20X 


1X 


3X 10X 


20X 


Electromechanical: 

Inverse 2 22.0 

Very inverse 4.15 

Extremely inverse 1.60 

Static: Inverse, very inverse, extremely inverse 6.42 


10.80 5.00 
4.15 2.90 
1 .60 1 .60 
1.5 .42 


3.66 

2.20 

1.60 

.31 


1.45 
.59 
.17 
.72 


0.65 0.32 
.58 .40 
.17 .17 
.174 .046 


0.24 
.25 
.17 
.033 


1 Coil current is expressed as a multiple of maximum coil range. 

2 The inverse induction-disk relay was not available with a range of 1 .5 to 12 A. The burdens given in that column are for relay range of 2 to 16 A on the 2.0-A 
tap. Comparison is still possible because the static-relay burden is constant regardless of the tap setting. 



362 



Of all the kinds of protection that can be provided by 
static relays, the use of zero-sequence relaying in high- 
voltage resistance-grounded systems is likely to give the 
most benefit to the mining industry. The increased sensi- 
tivity and decreased burden should remove many of the 
problems that were covered in chapters 9 and 10. The 
other major distribution protective-relaying concern is 
short circuit and overload, but induction-disk relays have 
already proved themselves to be effective and reliable 
here. For these relays to be replaced by static devices, 
lower cost and higher reliability must first be effectively 
demonstrated to the industry. 

A protection-relaying problem that has plagued the 
mining industry for many years is the occurrence of faults 
on dc trolley circuits. The situation can be divided into two 
main subjects (6). 

1. For short-circuit protection, the dc overcurrent relay 
(device 76) often works off a shunt. When using large 
locomotives, its pickup setting is on the order of 3,000 to 
4,000 A. Testing with such large values to trip electrome- 
chanical relays can be dangerous and the resulting accu- 
racy questionable. 

2. Because normal load currents can exceed 3,000 A, 
conventional electromechanical relaying cannot detect 
arcing or high-impedance faults, which may, for instance, 
lead to a mine fire. 

The best avenue for solving these problems appears to be 
through solid-state relaying. Static overcurrent relays af- 
ford safe, fast testing as required by Federal regulations 
(33), with inexpensive hand-held devices (6-7). For item 2 
above, many techniques are presently under demonstra- 
tion, and they were covered in chapter 9. 

When conventional molded-case circuit breakers are 
used for trailing-cable protection, the precision of short- 
circuit settings can always be questioned because of the 
mechanical nature of the trip elements. Overload adjust- 
ment is not available and, if possible, requires an ex- 
change of thermal elements. Mining-duty molded-case 
circuit breakers are available with static trip elements for 
both short circuit and overload. Here the conventional 
thermal-magnetic trip elements are replaced by CT's and 
solid-state circuitry. The CT's proportionately reduce the 
line currents to a level that can be used as input to the 
solid-state circuitry. Breaker tripping is initiated when a 
low-power flux-transfer shunt trip or UVR is activated 
from the output of the solid-state circuitry. The overload 
rating can be altered by simply changing a small rating 
plug on the front of the breaker. The instantaneous trip 
range is specified as a multiple of the overload-current 
setting. Operation of the static trip elements has shown 
their increased accuracy, reliability, and repeatability over 
their mechanical counterparts. 

It is interesting to note that in the United Kingdom 
static relays are used extensively for overload, short- 
circuit, and ground-fault protection of distribution and 
utilization equipment in mines (28). Two of the techniques 
used there deserve presentation: the sensitive earth- 
leakage system and the phase-sensitive short-circuit sys- 
tem. Recent research under Bureau of Mines funding has 
been investigating the adoption of similar systems to U.S. 
mine power systems (23-25). 



Sensitive Earth-Leakage System 

In an attempt to reduce the dangers of incendive 
arcing from damaged trailing cables, the U.K. National 
Coal Board developed the sensitive earth-leakage system 
(SEL) for ground-fault protection (32). The system also 
substantially reduces the chance of an electrocution by 
limiting ground-fault currents to extremely low values 
(27). 

A simplified diagram of the SEL system is shown in 
figure 14.31. The neutral-grounding impedance limits the 
maximum ground-fault current to 750 mA. Fault detection 
is by the zero-sequence scheme (see chapter 9), but a 
solid-state amplifier increases sensitivity to as low as 90 
mA versus 4 to 6 A with typical electromechanical devices 
on low-voltage systems in the United States. Currents 
below 90 mA are allowed to flow continuously, but cur- 
rents up to 750 mA are permitted for about 0.02 s. To 
provide selectivity, a 0.4-s time delay is introduced at the 



Core-balance 



Transformer 




Figure 14.31.— Simplified sketch of the SEL system. 



363 



main circuit breaker. The power factor of the neutral- 
grounding (earthing) impedance is normally specified 
from 0.65 to 0.75 to avoid limiting the current to a level 
that cannot be detected because of system capacitance. 
The circuit is simply a solid-state amplified zero-sequence 
relay. Because of the high sensitivity, grounded shields are 
placed over the CT to reduce electromagnetic interference 
that can cause nuisance tripping. 

To prevent the circuit from being reset until a ground 
fault is cleared, the SEL system has an auxiliary circuit 
connected to a second winding on the CT. Upon ground- 
fault pickup, an auxiliary contact is closed by the relay, 
which in turn causes the auxiliary CT winding to be 
energized. This induces a voltage on the other CT winding, 
which creates a lockout. 

The other ground-fault method in use in the United 
Kingdom is the multipoint SEL. Here a false-neutral 
transformer, which is impedance grounded, replaces the 
zero-sequence transformer. The source-transformer sec- 
ondary is isolated from ground across a spark gap (fig. 
14.32). When a ground-fault occurs, a potential is devel- 
oped across the wye-connected impedances (false-neutral 
transformer), and current flows through the grounding 
impedance. The voltage developed across the impedance is 
amplified, causing the relay to pick up. As with the 
preceding system, an auxiliary changeover contact pro- 
vides lockout until the fault is cleared. 

The multipoint system, however, has several disad- 
vantages: the technique is indiscriminate and limits the 
number of units that can be utilized at a gate-end box 
(utilization center). The maximum number of units for a 
550-V system is 37 and for a 1,100-V system, 18. Ground- 
fault current is again limited to 750 mA, but this time all 
relays see the fault current produced and will pick up. This 



Transformer 



Circuit 
breaker 




Earth fault relay 
contact in pilot circuit also) 



Auxiliary changeover 
switch operated 
by main contactor 




False- 
neutral 
transformer 
or wye-connected 

impedance 



Electronic 
amplifier 



Secondary supplying 
amplifier and earth fault relay 



EX 



Secondary supplying 
lockout circuit 




Figure 14.32.— Simplified sketch of the multipoint SEL 
system. 



is a definite drawback, even though the unfaulted units 
may be reset at once and the faulted unit will be locked 
out. In the United Kingdom, automatic circuit-breaker 
resetting is allowed on this system to restore operation to 
the unaffected portion. 

The sensitivity of the multipoint system is excellent: 
pickup is about 3 and 6 mA on 550- and 1,100-V systems, 
respectively. The high sensitivity is desirable in the case of 
two simultaneous ground faults on separate lines (27). 
Here, with a motor at full load, only 50% of system voltage 
is available to drive the ground-fault current. 

Phase-Sensitive Short-Circuit Protection 

Because both motor starting and fault conditions 
result in large current flows of comparable magnitude, it 
is often difficult to adjust standard relays to differentiate 
between the two and to provide interruption only when a 
fault occurs. Phase-sensitive relaying is able to distin- 
guish between these two conditions by sensing the phase 
angle between the current and voltage. It utilizes the fact 
that the induction-motor power factor at starting is ap- 
proximately 0.5, while typical faults have power factors of 
about 0.9 (14). Actual in-mine use of the method has been 
very promising, revealing that nuisance tripping on motor 
starting can be eliminated, while short-circuit trip set- 
tings can be reduced to half the value of those required for 
standard instantaneous relays (28). Two techniques can be 
used to provide the protection: a diode bridge and an 
electronic comparator. 

The circuit shown in figure 14.33 is a simplified dia- 
gram of one phase of a three-phase phase-sensitive system 
employing a switching diode bridge. In this circuit, the CT 
and diodes act as a standard full- wave bridge-rectifier circuit 
to create the voltage across the bridge (V b ). The PT second- 
ary may be modeled as the secondary of an ideal transformer 
in series with a resistor as shown in figure 14.34. The voltage 
in the secondary of the PT, V v , will be a sinusoid, and the 
resultant voltage across the transformer secondary resis- 
tances V r is V b + V v . Current flow in this resistance will be 
proportional to V r , as will the resultant current flowing 
through the CT and the load resistance R L . The trip signal 
will then be the voltage appearing across R L due to the 
current flowing through it. Thus, the trip signal ultimately 
is proportional to V v + V b . 

If the voltage and current are in phase, V v and V b are 
in phase, and they add in phase to produce V trip . If V b 
increases because of an increase in current flow, the peak 
value of V trip will increase. The instantaneous-relay unit 
connected across R L could be adjusted so tripping would 
occur when V trip reaches a certain magnitude that corre- 
sponds to a short-circuit condition on the system. If the 
voltage and current are out of phase by 60° (0.5 pf), 
relating to motor starting, V trip is less than for the 
in-phase situation. Even if the current were increased 
significantly (as in motor starting), V trip would still be 
lower than the voltage produced by a system fault, and the 
instantaneous unit would not trip. 

A block diagram of an electronic-comparator circuit 
that could also be used to provide this protection is shown 
in figure 14.35. The voltage and current comparators will 
each output a "1" (high logic level) when their respective 
inputs are above a specific threshold level. The AND gate 



364 




Figure 14.33.— Diode-bridge phase-sensitive protection. 




Figure 14.34.— Equivalent model of figure 14.33. 



from PT 



Iin 
from CT 




Figure 14.35.— Electronic-comparator method of phase- 
sensitive protection. 



will then output a "1" as long as the outputs of both the 
current and voltage comparators remain high. This AND- 
gate output signal is then integrated by the integrator 
circuit as shown, causing V trip to reach a level that would 



operate a circuit-interrupting device if the AND-gate 
output remained high too long. The integrator can be 
designed with enough leakage so V trip will not reach the 
tripping level during short-time intermittent overvoltages 
or overcurrents. 

If the voltage and current are in phase and a high- 
current magnitude exists (a fault condition), both V and I 
are above the comparator threshold levels, causing the 
comparator outputs to be high during the voltage and 
current peaks. Moreover, the voltage and current are in 
phase so the AND-gate output is high during these peaks 
as well. The AND-gate output is integrated, and V trip 
increases to the level where tripping occurs. It can be seen 
that normal current levels, even if they are in phase with 
the voltage, would not cause tripping because they would 



never cause the current comparator output to reach a high 
state. 

During motor starting, I and V are still above the 
comparator threshold levels, but they are out of phase, 
resulting in a narrower pulse at the output of the AND 
gate. Although the AND-gate output is integrated, the 
narrow pulse width allows the integrator to return V trip to 
zero before the next pulse occurs. This prevents short- 
circuit tripping for a motor starting condition on the 
system. Indeed, a separate relaying circuit would still be 
necessary for overload protection. 

This system protection scheme, which can be imple- 
mented either electromechanically or electronically, prom- 
ises to contribute toward meeting both of the conflicting 
objectives of coordination: protection and selectivity. Tests 
have shown that electronic phase-sensitive protection 
tripped at much lower levels and in shorter times than 
standard short-circuit devices. While typical instanta- 
neous relays may need a pickup setting of 7 to 10 times 
normal full-load current to prevent nuisance tripping on 
motor start, it has been found that phase-sensitive relay- 
ing eliminates spurious tripping during motor start, even 
when the pickup is as low as 3 times full-load current (14). 



SOLID-STATE RELAYS IN THE FUTURE 

The future development of static relays is indicated in 
many theoretical papers on the subject. Basically, new 
devices can be divided into three groups: continuous, 
digital-controlled continuous, and digital. Continuous re- 
lays are the types in which the power system is sensed 
continuously and relaying becomes activated from the 
sustained existence of a malfunction. A digital-controlled 
continuous technique has been illustrated in figure 14.35, 
and figure 14.36 relates its use in timed-overcurrent 
relaying. The full-wave rectifier and CT continuously 
sense power-system operation, and here the particular 
relaying characteristics are provided by the function gen- 
erator with time delay provided by the linear-ramp gener- 
ator. The pickup-level and ramp-level detectors essentially 
function as digital control devices since they operate in a 
go or no-go manner (19). 

Digital relays rely on discrete sampling of current 
and/or voltage waveforms, directing this data through an 
algorithm and decision process, which ultimately decides 
whether a relay is to be actuated. The ultimate in digital 
relays of the future will probably use microprocessors 
distributed throughout the power system to analyze cir- 
cuit conditions on a real-time basis. Each microprocessor 
could be responsible for making decisions for its system 
portion with control relinquished when demanded by a 
larger computer or when power-system conditions exist 
that are beyond the analytical capability of the particular 
microprocessor in question. 



SUMMARY 

The solid-state applications discussed in this chapter 
provide a viable alternative to their electromechanical 
counterparts. Indeed, most of them are decidedly superior 
in carrying out conventional functions, and in some cases 
solid-state relays exhibit characteristics and can perform 
functions that are not available with other devices. A 
major drawback with static belt starters and relays is their 



365 



Power- 
Q — I supply CT 



c 



c 



Rectifier 

and 
voltage 
regulator 



Signal 
source CT 



■ +15.0 

■ +5.1 
. COM 

■ -5.1 

■ -15.0 

Pickup 
setting 



Time- Time- 
dial dial 
switch vernier 



Time 
calibration 



Full- 
wave 
rectifier 



Current- 
tap 
resistors 



Instantaneous 

overcurrent 

unit 



Function 
generator 




Time- 
dial 
resistors 



Linear- 
-•] ramp 
generator 




Pickup- 
level 
detector 



Ramp 

level 

detector 



Output 
circuit 



Pickup 
calibration 



Telephone 
relay 



Time 

target 

unit 

Figure 14.36.— Digital-controlled continuous static relay used for timed overcurrent. 



cost, and the major factor in the cost breakdown is reli- 
ability. Belt starters have proved their desirability 
through a decade of in-mine experience, whereas the 
reliability of static relays has not been well documented to 
date, especially in mining. This problem is due to their 
recent introduction and is compounded by the fact that 
older models have been replaced by newer designs before 
life test data could be compiled (33). Static relays presently 
occupy only a small portion of the protective-relaying 
market. Currently, the main benefits they provide for the 
mining industry are decreased burden problems, increased 
sensitivity for zero-sequence relaying in resistance- 
grounded systems, and their use as solid-state tripping 
elements in molded-case circuit breakers. When their 
reliability is more fully documented and is expressed in 
terms of cost effectiveness, they may well become more 
competitive with electromechanical units and expand 
their role to include overload and short-circuit protection. 



REFERENCES 

1. Andreiev, N. Power Relays Solid State. Vs. Elec- 
tromechanical. Control Eng., v. 20, Jan. 1973. 

2. Berman, S. W. Solid-State Reduced Voltage Starters for AC 
Induction Motors. Pres. at IEEE-IAS, Portland, OR, Chapter, 
Dec. 10, 1974. 

3. Boeniger, J. Solid State Overcurrent-Time Relays. Brown 
Boveri Rev., v. 58, July 1971. 



4. Bucheridge, R. M. Solid-State Controls for Mining Applica- 
tions. Pres. at 1975 Annu. Meet. IEEE-IAS, Atlanta, GA, 
Sept.-Oct. 1975. 

5. Bullock, S. A. Line Power Solid-State Motor Starter. Line 
Power Manufacturing Corp., Bristol, VA, undated. 

6. Burr, J. F. Solid-State Overcurrent Relays for the Protection 
of Trolley Distribution Systems in Underground Mines. Paper in 
Conference Record- IAS 11th Annual Meeting (Chicago, IL, Oct. 
1976). IEEE, 1976. 

7. Chumakov, W. V. Protective Relaying for Mining Applica- 
tions. Pres. at 3d Annu. West. Protective Relay Conf., Spokane, 
WA, Oct. 18-21, 1976; available from BBC, Allentown, PA. 

8. Collins, H. W. Try Solid State Relays for Industrial Control 
Systems. Automation, v. 19, Mar. 1972. 

9. Dewan, S. B., and A. Straughen. Power Semiconductor Cir- 
cuits. Wiley, 1975. 

10. Downs, C. L., and W. V. Chumekov. Solid-State Relaying 
Applied to High-Voltage Distribution in Mining. Paper in Con- 
ference Record -IEEE Mining Industry Technical Conference 
(Pittsburgh, PA, June 1979). IEEE, 1979. 

11. Einvall, C. H., and J. K. Lander. A Static Directional Over- 
current Relay With Unique Characteristics and Its Applications. 
Pres. at IEEE-PES Winter Meet., 1976. 

12. Fitzgerald, A. E., D. E. Higginbotham, and A. Grabel. Basic 
Electrical Engineering. McGraw-Hill, 4th ed., 1975. 

13. Gehmlich, D. K., and S. B. Hammond. Electromechanical 
Systems. McGraw-Hill, 1967. 

14. Gray, G. W. Phase Sensitive Short-Circuit Protection. Min. 
Technol, v. 60, Apr. 1978. 

15. Institute of Electrical and Electronics Engineers (New 
York). IEEE Standard Dictionary of Electrical and Electronics 
Terms. IEEE Stand. 100-1984. 



366 



16. Johnson, P. Solid-State Relays-A Guide to Their Design 
and Applications. Electron. Des. News, v. 18, Oct. 5, 1973. 

17. Konrad, C. E., and B. Harbage. Solid State Control of Trac- 
tion Motor Devices for Large Battery-Powered Vehicles. Pres. at 
L974 Annu. Meet. IEEE-IAS, Pittsburgh, PA, Oct. 1974. 

18. Lyons, R. E. Electromechanical Solid State. Mech. Eng., v. 
94, Oct. 1972. 

19. Manley, R. C. A New Solid State Time Overcurrent Relay 
for Industrial and Commercial Power Systems. Paper in Proc. 1975 
Annu. Meet. IEEE-IAS. 

20. Matuszak, R. A. Controlled Frequency -The Brushless 
Electric "Steam Engine." Min/ Eng, v. 30, Feb. 1978. 

21. Millermaster, R. A. Harwood's Control of Electric Motors. 
Wiley-Interscience, 4th ed., 1980. 

22. Mohr, J. C, and P. Farinas. Solid State Overcurrent Relays 
Bow In. Power, v. 116, Mar. 1972. 

23. Morley, L. A., T. Novak, and F. C. Trutt. Electrical-Shock 
Prevention (contract J01 13009, PA State Univ.). Volume 
I -Protection of Maintenance Personnel. BuMines OFR 177(l)-83, 
1982; NTIS PB 84-102946. 

24. Morley, L. A., F. C. Trutt, and T. Novak. Sensitive Ground- 
Fault Protection for Mines. Phase I. Alternating-Current Utiliza- 
tion (contract J0134025, PA State Univ.). BuMines OFR 26-85, 
1984; NTIS PB 85-185767. 

25. Morley, L. A., F. C. Trutt, and D. J. Rufft. Electrical Shock 
Prevention (contract J01 13009, PA State Univ.). Volume 
II -Ground-Fault Interrupting Devices. BuMines OFR 177(2)-83, 
1982; NTIS PB 84-102953. 

26. Osborn, G. Solid State Catches Up in Relay Race. Electron. 
& Power, v. 18, Aug/Sept. 1972. 

27. Pearson, F. K., and G. B. Scott. Earth Leakage Protection 
for Coal Mine Electrical Supplies. Min. Technol., v. 53, Mar. 1971. 

28. Scott, G. B. Ground-Fault and Short-Circuit Protection in 
Underground Coal Mines- UK Practice. Paper in Conference 



Record -IEEE Mining Industry Technical Conference (Pittsburgh, 
PA, June 1979). IEEE, 1979. 

29. Sedgwick, A. Are Solid State Relays Really Necessary? 
Electron. Eng., v. 44, May 1972. 

30. Sheffer, K., D. L. Graham, and R. E. Obenhaus. Application 
of Inherent Thermal Protection to Industrial Motor Systems. Pres. 
at 1972 Annu. Meet. IEEE-IAS, Philadelphia, PA, Oct. 1972; 
available from 

31. Smith, H. D. Application of Static Starters in Underground 
Coal Mines. Pres. at 1974 Annu. Meet. IEEE-IAS, Pittsburgh, PA, 
1974; available from 

32. U.K. National Coal Board (London). Colliery Electrician. 
Butler and Tenner Ltd., 1976. 

33. U.S. Mine Enforcement and Safety Administration (now 
U.S. Mine Safety and Health Administration). Guidelines for 
Evaluation of Metered Relays and Solid-State D.C. Trolley and 
Trolley Feeder Wires. Jan. 22, 1976. 

34. Waldron, J. E. Innovations in Solid-State Protective Relays 
for Industrial and Commercial Power Systems. Pres. at IEEE Ind. 
and Commer. Power Syst. Tech. Conf., 1975. 

35. Warner, E. M. Status Report on Solid State Controls for 
Underground Mining Machines. Min. Congr. J., v. 59, Mar. 1973. 

36. Warrington, A. R. van C. Protective Relays, Their Theory 
and Practice. Wiley, v. 1, 1962. 

37. Protective Relays, Their Theory and Practice. 

Wiley, v. 2, 1977. 

38. Westinghouse Electric Corp., Relay-Instrument Div. 
(Newark, NJ). Applied Protective Relaying. Silent Sentinels Publ., 
1976. 

39. Wetter, C. R. Review of AC/DC Solid-State Controls: Past, 
Present, and Future Usage. Min. Congr. J., v. 63, Apr. 1977. 



367 



CHAPTER 15.— BATTERIES AND BATTERY CHARGING 



Although the storage battery had a variety of appli- 
cations in the 1800's, its successful use for traction pur- 
poses was not achieved until the turn of the 20th century. 
Early mining batteries were used to power gathering 
locomotives, and to a certain extent they replaced mules in 
nongassy mines where open lights were used. From these 
beginnings, the battery-powered vehicle has gradually 
increased in popularity to become an important part of 
many underground coal mines, both for rail and off-track 
haulage. 

The first haulage applications had little concern for 
safety, but the storage-battery locomotive was soon recog- 
nized as having inherent safety advantages over trolley 
locomotives and cable-reel locomotives. As early as 1919, 
Appleton noted that batteries reduced the chance of fire 
from short circuits and arcing (I). 1 Ilsley (15) said of the 
battery locomotive: 

That its energy is self-contained and limited to 
the immediate zone of the locomotive is a safety 
factor of great importance. In the trolley type of 
equipment one necessarily uses the track return, 
and the danger zone from the return current may 
extend through the mine. Poor bonding or no 
bonding may force the return current back toward 
the face. A storage battery locomotive does not 
use or need the dangerous overhead trolley with 
its constant shock menace and fire hazard, and 
with the possibility of trolley or feeder circuits 
becoming a factor in the ignition of gas or coal- 
dust. 

Storage batteries have had a relatively good underground 
safety record. 

However, early battery locomotives had open control- 
lers, open motors, weak battery covers, crude exposed 
wiring, and battery jars that were prone to breakage. To 
help correct these problems, the Bureau of Mines in 1919 
issued Schedule 15, which set standards for permissible 
battery-locomotive equipment (30). Even with this regula- 
tion in force, several incidents involving battery gathering 
locomotives occurred that raised further questions con- 
cerning battery safety underground. For instance, a major 
mine explosion in Everettville, WV, claimed the lives of 97 
miners, and the resulting investigation traced the ignition 
to a battery locomotive (5). A report by Owings (21) 
contained a description of another accident where the 
mine operators felt strongly that the cable-reel locomotive 
was much safer than battery power in gaseous mines. 
However, Owings examined the facts pertaining to both 
incidents and concluded that battery locomotives were 
relatively safer than the other available gathering- 
locomotive types. The Everettville explosion was created 
by a nonpermissible locomotive and could just as easily 
have occurred from the use of other nonpermissible equip- 
ment. The second incident was later discovered to have 
resulted from a faulty cell that had accidentally been 
reversed, causing the cell to emit hydrogen, which ex- 
ploded and dislodged the box cover. Owings concluded that 
the occurrence was very rare and that the risk of fire was 
greater when trailing cables were involved. 



1 Italicized numbers in parentheses refer to items in the list of references 
at the end of this chapter. 



The shuttle car, introduced in 1938, was initially 
battery powered, but cable-reel shuttle cars soon became 
much more popular than the battery cars, probably be- 
cause of the low-capacity per-unit weight of the batteries 
then available. In the United Kingdom, trackless, battery- 
powered vehicles were widely believed to be unsafe. As a 
result of explosions caused by storage batteries at Weet- 
slade and Eppleton collieries, battery-powered trackless 
vehicles were prohibited from nearly all British mines and 
the use of battery locomotives was restricted to types 
approved by the Government (2). In the United States, 
however, permissible battery-powered equipment is al- 
lowed inby the last open crosscut. 

Although batteries are not currently used to power 
shuttle cars, they have a variety of other uses in coal 
mines, including the provision of cap lamp power. There 
has been a recent increase in popularity of battery- 
powered vehicles because the efficiency of batteries has 
been greatly improved. One prominent manufacturer re- 
ported a 70% increase in ampere-hours per cubic foot and 
a 39% increase in watthours per pound over earlier meth- 
ods (12). In addition, improvements in plate and grid 
design have increased the average service life of motive 
power batteries. The trend in underground mining is 
toward the use of conveyor belts for coal transport and rail 
systems for personnel and supply movement, but mines 
that have ac face equipment and dc trolley lines can be 
subject to nuisance tripping of ac circuit breakers from 
stray dc ground currents. These currents, which often 
result from poor track bonding, ineffective trolley-line 
insulators, or inadequate dc-to-ac ground-system isolation, 
can be eliminated when batteries are used to power the 
locomotives. 

Batteries are also used to power articulated ram- 
dump haulers, tractor-trailer units, or scoops (front-end 
loader tractor units). These battery-powered face-haulage 
units eliminate problems involved with shuttle car trail- 
ing cables, and in small conventional mines, the operator 
can be spared the cost of a loading machine and a separate 
machine for cleanup and supply haulage. For moderate- 
to-large production operations, the extreme mobility of the 
tractor-trailers and scoops has made them invaluable 
ancillary equipment for cleanup and supply in practically 
all longwall and many continuous operations. This chap- 
ter will examine the application of batteries to power such 
equipment. 

Some inherent hazards in battery use remain: batter- 
ies emit hydrogen, an explosive gas, while charging; 
batteries and battery chargers are capable of delivering a 
fatal electric shock; and batteries can be a potential fire 
hazard. A number of less catastrophic hazards may also be 
encountered, ranging from acid burns from spilled electro- 
lyte to pinched fingers from careless handling. 



BASIC BATTERY AND BATTERY-CHARGING 
THEORY 

A storage battery can be defined as a battery in which 
the electrochemical action is reversible (7); that is, after an 
output of electrical current (discharge), the battery can be 
returned to the original state (recharged) by passing 
current through it in the opposite direction. The basic unit 



368 



of the battery is the cell, which simply consists of positive 
plates, negative plates, and electrolyte. One or more cells 
are connected together to form the battery. The connection 
is usually in series, although parallel and series-parallel 
combinations are sometimes used. The battery voltage, 
often given as an open-circuit value, is the sum of the 
series cell voltages. The battery capacity is commonly 
expressed in ampere-hours or kilowatthours and is mainly 
dependent upon the plate size (surface area). 

Two types of storage batteries have been employed in 
underground traction, alkaline and acid. The nickel-iron 
or Edison cell is an alkaline cell because of the type of 
electrolyte used. The plates for this battery are con- 
structed of nickel oxide and iron, immersed in an electro- 
lyte of potassium hydroxide and lithium hydroxide. Edison 
batteries were once popular in the mining industry be- 
cause of their high reliability and low-maintenance char- 
acteristics, but lead-acid batteries have now entirely re- 
placed the Edison type, as a result of their high energy per 
unit volume and high power capability. 

The basic lead-acid cell utilizes a lead peroxide (Pb0 2 ) 
positive plate and a sponge lead (Pb) negative plate. These 
plates are suspended in a solution of dilute sulfuric acid 
(H 2 S0 4 ) (fig. 15.1) (18). When a circuit is completed be- 
tween the positive and negative plates, the following 
reaction occurs: 



Pb0 2 + Pb + 2H 2 S0 4 discharge 
charge 



2PbS0 4 + 2H 2 + 2e - 
(15.1) 



The battery must be recharged when a large portion of the 
Pb0 2 and Pb is in the form of PbS0 4 and H 2 0. Charging a 
battery consists of supplying electricity to drive the reac- 
tion as shown. When fully charged, the positive-plate 
active material is all lead peroxide, that of the negative 
plate is all sponge lead, and the specific gravity of the 
electrolyte is at a maximum (7). 

The lead-acid cell has the highest voltage of any 
commercial battery cell. The nominal voltage is generally 



referred to as 2.0 V, but the actual value varies depending 
upon the electrolyte specific gravity and whether the cell 
is being charged or discharged. The open-circuit voltage is 
around 2.12 V per cell, with a full-charge specific gravity 
of about 1.280, which is a common level for motive storage 
batteries (7, 18). However, because of the effective internal 
resistance, the cell voltage decreases as soon as discharge 
commences. As illustrated in figure 15.2 the voltage 
continues to decrease with discharge, and the rate of 
voltage decrease is connected to the discharge current rate 
(7). The final voltage is the point where the battery is no 
longer effective for its application. Because this value also 
changes with the discharge rate, a typical final voltage of 
1.75 V is often assumed. A standard discharge time is 
taken as 8 h. When the discharged battery is placed on 
charge, cell voltage immediately increases and continues 
to rise as the charging processes (fig. 15.3) (18). 

The number of times a lead-acid battery can be 
recharged is a function of the discharge level of the battery 
during its working cycle, the method used to charge the 
battery, and the quality of battery maintenance. Each 
parameter is independent of the others. The effect of the 
discharge level on N, the number of times a battery can be 
recharged is given by 



N = 



(15.2) 



D x 



where k and x are constants of the particular battery, and 
D is the percent of total battery energy removed during a 
typical discharge. The equation states that the theoretical 
number of times a battery can be recharged is inversely 
proportional to the discharge level raised to the power x. 
Battery manufacturers recommend that lead-acid batter- 
ies should be recharged when they reach 80% of their 
capacity. Because it is difficult to determine the extent of 
battery discharge until the battery is dead, a battery 
should be recharged whenever the machine it is powering 
begins to show signs of sluggishness. The discharge should 



Lead 
peroxide 




PARTIALLY CHARGED OR 
PARTIALLY DISCHARGED 



Sponge 
lead 



Lead 
peroxide 
and lead 

sulfate 



Full- 
strength 
electrolyte 
sulfuric acid 
and water 




DISCHARGED 

+ 



Lead 
sulfate ' 



Sponge 

lead 

and lead 

sulfate 



Medium- 
strength 
electrolyte 
sulfuric acid 
and water 



•' 



Lead 
sulfate 



Weak 
electrolyte 

sulfuric 
acid and 

water 



Figure 15.1.— Composition of lead-acid storage battery in various states of charge. 



369 



not reach the point of cell exhaustion or where voltage 
drops below a useful value (7). 

About 110% of the ampere-hours discharged must be 
returned to fully charge a lead-acid battery. The rate at 
which charge is restored is an important consideration in 
attaining the maximum number of charge cycles and 
maximum life of the battery. Generally, any current level 
is acceptable, provided that temperatures above 115°F are 
not produced (125°F for short periods) and excessive gas- 
sing does not occur (18). Manufacturers may publish a 
normal or finish rate; this is also the current level at 
which the battery can be safely charged any time charging 
is required. The charge cycle is usually 8 h long. 

Most of the electricity supplied to a cell being charged 
is used to transform water and lead sulfate into sulfuric 
acid, lead, and lead peroxide, but some of the current 
causes electrolysis, breaking the water down into its 
constituents, 



2H 2 - 2H 2 + 2 , 



(15.3) 



which is called gassing. The rate of gassing increases 
dramatically at a cell voltage of 2.37 V when increasing 





<LA 




1 1 


1 1 ' 1 ' 1 


1 1 


I 




2.0 










- 


> 












- 


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(9 


i q 


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< 




\ 






^v^rate 




1- 




\ 










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-\ 




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- 


> 






\ 


\ \5h >^ 






_1 


1 .8 


— 


\ 






^w — 


_l 
HI 






V.h 


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o 




— 


i 1-h 

* rate 
1 


i *." Final 


volts 






1.7 




\ ,' 




— 








1 • 














\,' 






- 




1 6 




1 


. I.I.I 


, 1 


i 



4 6 8 

TIME, h 



10 



12 



Figure 15.2.— Voltage per cell of a typical lead-acid battery 
with varying continuous rates of discharge. 



UJ 

cr 
a: 

CJ 

o 

z 
o 

< 

X 

o 



140 
120 

100 
80 
60 

40 
20 



- 


1 1 1 


1 1 1 


1 


- 


Charging current^ 




— 




Cell voltage^ ^/\ 


- 


- 


1 1 1 


1 1 1 


1 



1.120 



2.7 
2.6 

2.5 




2 4 uj 
< 

23 b 

£..0 Q 
> 




2.2 j 

UJ 

o 
2.1 




2.0 




120 




too 




80 o" 




ul 

60 oc 
< 

X 

40 lj 
cc 


u_ 


20 



-,100 °„ 
-1 80 lu 



Figure 15.3.— Typical charging process of cell from 18-cell, 725-Ah battery. Ambient temperature, 77°F. Specific gravity 
temperature adjusted. 



370 



quantities of current become available for electrolysis 
because of the higher state of charge in the cell. 

A certain amount of gassing is a necessary conse- 
quence of a good charge, which explains why water must 
periodically be added to batteries, but excessive gassing or 
overcharging causes damage to the plates, excessive water 
consumption, and excessive hydrogen emission. For this 
reason, the rate of charging current must be controlled as 
the battery charges. The large amounts of H 2 and 2 
released during excessive gassing cannot be detected by 
mine personnel. However, the H 2 and 2 are sometimes 
accompanied by amounts of H 2 S0 4 released into the mine 
atmosphere, which is easily identified by smell. 

Undercharging a battery is also harmful, even if it is 
only practiced occasionally. Insufficient charging leads to 
a gradual sulfation of the negative plates, which eventu- 
ally causes a reduction in battery capacity and life (7). A 
similar situation can occur when batteries are left stand- 
ing in an uncharged state for long periods (18). 

BATTERY MAINTENANCE 

Proper battery maintenance has a very significant 
influence on battery life. Battery manufacturers normally 
include a recommended maintenance program with their 
batteries, based on specific gravity levels and equalizing 
schedules. Equalizing is the process by which all the cells 
in a battery are brought to the same voltage. 

In lead-acid batteries, the electrolyte specific gravity 
is a function of the state of battery charge. Consequently, 
a plot of electrolyte specific gravity versus discharge depth 
for a particular battery is important. All lead-acid batter- 
ies require periodic equalization, but excessive equaliza- 
tion can cause unnecessary battery deterioration. 

The following guidelines can be used to develop a good 
battery maintenance program, but see also references 3-4, 
6, 8-9, and 14. There are three groups of activities in the 
program: those that should be performed daily or during 
each charge period, those that should be performed 
weekly, and those to be performed approximately once 
every 3 months. Accurate records should be kept of all 
maintenance activities for each battery as these are a 
convenient way to monitor individual battery perfor- 
mance. Deteriorating battery conditions can thus be de- 
tected before the battery becomes a safety hazard or the 
source of costly downtime. 

Daily battery maintenance activities should include 
the monitoring of one battery cell, which is called the pilot 
cell. Any battery cell can be used as the pilot. The 
following cell characteristics should be recorded during 
each charge: specific gravity before and after charging, 
electrolyte temperature before and after charging, and the 
water level in the cell. If any of the pilot parameters falls 
outside those specified as acceptable by the manufacturer, 
all cells should be checked and corrective action should be 
taken. Other daily maintenance activities should include 
checking the battery for physical defects such as cracked 
cell plugs and ensuring that the charger output voltage is 
correct. 

The weekly maintenance program includes checking 
all battery cells for proper water level (the water consump- 
tion of a good battery is generally equally distributed 
among the individual cells) and routine cleaning of the 
battery tops. Every 3 months, it is good practice to take a 
complete set of cell voltage and specific-gravity readings at 
the end of an equalizing charge to ensure that these 
parameters meet manufacturer specifications. 



CHARGERS 

The current supplied to a cell must be dc. It would be 
possible to obtain this current from a trolley-distribution 
system, but this source has a serious drawback: it is 
difficult to obtain the precise current requirements for 
charging. As a result, most mine batteries are charged 
from the ac distribution system, using a transformer- 
rectifier combination. Mercury arc rectifiers were used 
previously (10), but selenium or silicon rectifiers are now 
universally employed, and silicon diodes are considered to 
be the industry standard. Transformers generally have 
isolated secondaries and are most often three phase, 
although single phase can be used. Rectifiers are usually 
in a full-wave bridge. 

Several methods can be used to control the rate of 
charge. Most are designed to initiate the charge at a fairly 
high current or starting rate, commonly 20 to 25 A per 
100-Ah capacity. The level is then tapered off to the finish 
rate as the battery charge is restored, for example, 4 to 5 
A per 100 Ah, as shown in figure 15.3. Some chargers 
reduce the starting rate to the finish rate in one step when 
the charge is about 80% complete, but control devices that 
taper the charge rate are more common in mining. Taper 
chargers are either active or passive, and in both types the 
charge is usually stopped automatically when full charge 
is reached (7). 

The most popular passive system for taper charging 
employs some value of ballast resistance placed in series 
with the battery (32). This resistance limits the initial 
charge current and gives a relatively flat current-versus- 
time curve throughout the charging cycle. The advantage 
of this method (sometimes termed modified constant poten- 
tial) is its simplicity. Its disadvantages include the fact 
that the ballast resistor dissipates a rather large amount 
of energy and the charge rate does not coincide with that 
considered optimal for lead-acid batteries. A variation of 
the above system uses a timer to switch additional resis- 
tance in series with the battery, thereby reducing the rate 
of charge at some point in the charge cycle. 

A characteristic common to all active systems is 
control of the charging current by a feedback system that 
samples the battery voltage during charging. The simplest 
active system in taper chargers consists of a voltage- 
controlled relay that switches additional resistance in 
series with the battery at a cell voltage of 2.37 V. This and 
all active systems must have different voltage thresholds 
for batteries with different numbers of cells. 

Another active method, shown in simplified form in 
figure 15.4, utilizes a saturable reactor type of voltage 
transformer to feed the rectifier. The saturation level of 
the transformer core is controlled by a dc applied to a 
winding on the core so that the secondary voltage is 
regulated by voltage-controlled current feedback from the 
battery. Output regulation usually begins at a battery 
voltage of 2.37 V per cell and is varied to zero current at 
the end of the charge cycle (9). 

A relatively new method uses thyristors to replace 
half of the silicon diodes in the full-wave bridge, as shown 
in figure 15.5. The firing angle is determined by a feed- 
back circuit that senses battery voltage (17). The circuit 
given is for a single-phase charger. Three-phase chargers 
would have three SCR's in the full-wave bridge with three 
firing boards, but the control circuitry would be much like 
that for single phase. Although many manufacturers and 
users have had success with full-wave rectifier configura- 
tions, in some designs a phase reference for firing the 



371 



Transformer 



Core 



O-d 



Main 
contactor 




Saturable 
reactor 



Timer 



Figure 15.4.— Simplified schematic of saturable-reactor 
charger. 




I 

J Battery 

:=: being 
i charged 



m K h m 



r\ r\ r\ r\ 



Single-phase line voltage 



SCR bridge output voltage 
■Battery voltage 
Firing board output pulse 

Current output 



Figure 15.5.— Simplified schematic of single-phase thyristor 
charger. 

SCR's cannot be obtained with a full wave because of 
negligible ripple. As a result, a half- wave must be used 
and this gives less efficient rectification and sometimes 
produces a chopped waveform, which is not entirely suit- 
able for battery charging. 

The last active method employs a ferroresonant trans- 
former and bucking coils (13). For the conventional trans- 
former, as modeled in figure 15.6, any change in primary 
voltage produces a corresponding change in secondary 
voltage, which is maintained until saturation commences, 
the knee shown in the transformer magnetization curve of 
figure 15.7. This is an advantage for most power applica- 
tions since transformer operation approaches ideal when 
used in the linear portion of the curve (to the left of the 
knee). However, there are many instances where fluctua- 
tions in the secondary voltage due to normal primary- 
voltage changes are unwanted, and here the secondary 
voltage must be regulated, or remain reasonably constant 
under a specific load condition. A more constant but 
distorted secondary voltage could be obtained by driving 
the transformer into the saturation region (to the right of 
the knee in figure 15.7). Here, any change in the primary- 
voltage magnitude would cause only a small change in the 




Figure 15.6.— Two-winding transformer model. 



x 




V V 

P1 P2 



V V 

P3 P4 



MAGNETIZING FORCE 

Figure 15.7.— Representative transformer magnetization 
curve. 



core flux and thus a small change in induction in the 
secondary winding, but this operation is unwise because 
primary current becomes excessive. Nevertheless, regula- 
tion can be obtained with normal primary current if the 
transformer is modified so the secondary winding is ex- 
posed to a saturated core portion while the primary 
operates under unsaturated conditions. This is the basic 
principle behind a ferroresonant transformer. 

A model of a basic ferroresonant transformer is shown 
in figure 15.8 (13). The leakage block of magnetic material 
provides a shunt path to bypass part of the flux produced 
by the primary winding, with the air gaps limiting the 
amount of flux bypass. The flux not bypassed causes 
induction in the resonant winding. The inductance and 
capacitance of this winding are selected so that additional 
flux is produced that is in phase with the primary flux 
when there is no load on the secondary winding. This 
increases the flux in the right core portion to about the 
knee of the magnetization curve. As a result, the right 
transformer portion will operate under saturated core 
conditions while the primary sees an unsaturated core. 

In the basic ferroresonant transformer, an increase in 
primary voltage will still produce a slight increase in 
secondary and resonant winding voltages. lb offset any 
increase in secondary voltage, a compensating or bucking 
coil, consisting of a few turns wound directly over the 
primary, can be connected in series with the secondary 
winding but with opposing polarity (fig. 15.9). Any in- 
crease in primary voltage will produce a proportional 
change in V b , which will offset any increase in secondary 
voltage and produce a rather constant output voltage over 
a specified range of primary voltages. 



372 



Figure 15.10 shows the adaptation of a ferroresonant 
transformer to a battery charger. Although a single-phase 
unit is illustrated, the technique can also be used in 
three-phase units. Some chargers do not use the bucking 
coils for reasons discussed later in the chapter. 



-^ 



Primary flux 



J2 



Air 

gaps 



u; — i u^ 

i » Resonant <-rE ; r~l 
It winding ,PE \ =3 
T Secondary -^^Ir' 



j' k Secondary ~* i ( ^&[ J 






C j) V c 



Leakage 
block 



Secondary 
flux 



Figure 15.8.— Ferroresonant transformer model. 



v f 


f Primary^ 
| coil 3 


Ic_ 


± > 

TV C 


Resonant 3 
coil 5 



C^ Bucking | v 
C^ coil 



: Secondary 
coil 



iut Load 



Figure 15.9.— Ferroresonant transformer. 



Active charging systems have an advantage over pas- 
sive systems in that feedback techniques can be used to 
give more accurate control of the charge rate. The most 
popular chargers for mine vehicle batteries use saturable 
reactors, thyristors, or ferroresonant transformers. Charge 
termination can be achieved by a timer or by monitoring 
the cell voltage or its rate of change to determine when the 
battery is fully charged. 

Although the success or failure of the battery-powered 
mine transportation system is largely a function of the 
operator's ability to get maximum life from the batteries, 
it is also a function of the safety factor involved in battery 
usage. The basics of battery safety will therefore be 
discussed in some detail in the following sections. 



CHARGING STATIONS 

Special charging stations are required in under- 
ground mines to charge vehicle batteries. These stations 
must be designed and constructed to meet specific venti- 
lation requirements, but because mining methods and 
plans vary widely, it is extremely difficult to define a 
rigorous set of guidelines. Hence, this section presents the 
principles that underly charging-station construction and 
explains the requirements that must be met. 

The first problem that must be addressed at battery 
charging stations is dissipation of the gas produced by the 
charging operation. It has already been stated that hydro- 
gen gas is liberated at the close of the charge cycle. 
Although modern chargers are designed to prevent exces- 
sive gassing by automatically dropping the charge rate to 
a very low value when a specified cell voltage is reached, 
it is impossible to charge a battery properly without 



To ac 
power supply ^ 



Line 
contactor 



._ Timer 
T contacts 



s*\ Line 

( ) contactor 
V-^ coil 



6 



Timer 
motor 



Primary (£_ 



coils 



ac 
fuse 



HI DZ_Th 



Line 
contactor 



^-pv '-r - ^ ' 



Resonant 
coil 



Overvoltage 
suppressor 




Ammeter 



Figure 15.10.— Ferroresonant battery charger. 



373 



producing some gas. This gas is explosive and must be 
diluted to render it harmless in the mine atmosphere. The 
traditional method of achieving this is by forced ventila- 
tion of the charging room. The hydrogen concentration 
must be kept below its lower explosive level of 4%. Natu- 
rally, because of the catastrophic nature of underground 
explosions, hydrogen concentrations are closely controlled 
by Federal regulations. These limit the permissible con- 
centration of hydrogen in coal mine atmospheres to 0.8% 
by volume, which provides a safety factor of 5 (30 CFR 
75.301.5). This limit is analogous to a maximum allowable 
concentration of one-fifth of the lower explosive level of 
methane. 

In order to maintain these requirements, it must be 
possible to monitor or estimate the evolution of hydrogen 
in the battery. By assuming a typical charge characteris- 
tic, manufacturers have made it possible to estimate 
hydrogen evolution from the number of cells and rated 
ampere-hour capacity. Table 15.1 lists such formulas from 
three different manufacturers. It can be seen that the H 2 
evolution calculated from these equations is fairly consis- 
tent, ranging from 0.0024 to 0.0028 ft 3 per cell ampere- 
hour. These figures are applicable to the latter stages of 
the charge cycle at cell voltages of 2.37 V and greater. 
These estimated values have been confirmed by compari- 
son with actual concentrations and charging room airflow 
rates obtained by in-mine survey {19). 



_J 

Fresh-air , 



entry 



=> 



Charger 



2\ Fresh-air flow 




Corrugated 
metal 
siding 



*/////////////////, 




Scoop 
tractor 



140 ft 3 /min 
minimum 







i 




Roof 

support 

crib 








1 



•Blocking 



<^ 



Return-air 
entry 



Ventilation' 
opening 



L 



Y///////A 



/////////////////////////////////////// 



Figure 15.11.— Plan of underground charging station. 



Table 15.1.— Formulas to estimate hydrogen evolution 



Formula for ft 3 H 2 


H 2 evolution, 




Manufacturer liberated in last 


ftVhper 


References 


3 h of charge 1 


cell Ah 




C & D (number of cells) (Ah) (0.0024) 


0.0008 


3-4 


Exide (5) (Ah/100) (number of cells) 


.0008 


6 


(0.016) 






KW (number of cells) (Ah) (0.002948) .. 


.0007 


17-18 



For KW, last 4 h of charge. 



The second requirement for charging stations that is 
stipulated by Federal law is that all underground battery 
charging stations must be "housed in fireproof structures 
or areas." To render the station fireproof, it is common 
practice to line the charging area with corrugated metal 
siding so that all exposed coal (which has previously been 
rock-dusted) is covered. Concrete block is used as a lining 
in some instances but is more costly. 

The third Federal stipulation is that "air currents 
used to ventilate structures or areas enclosing electrical 
installations shall be coursed directly into the return." 
This implies that the charging station must have a sepa- 
rate split of fresh air. To comply with this ventilation 
requirement, stations are frequently located in unused 
crosscuts immediately adjacent to return-aircourse en- 
tries. A small opening is made in the ventilation blocking 
so that fresh air passes over the station and dumps 
immediately into the return. Figure 15.11 illustrates a 
charging station that meets Federal regulation. Many 
other configurations are possible. 

When each charging station is designed, it must be 
verified that the hydrogen concentration is below the 
maximum allowed. A simple but effective approach is to 
establish a worst case airflow quantity for each battery on 
charge, then substitute the safe condition by an actual 
airflow measurement of the charging station. This mea- 
surement can be made with an anemometer. 



From table 15.1 it can be seen that batteries in the 
final stages of charge evolve hydrogen to the following 
approximate level: 

(5) (Ah/100) (number of cells) (0.016) = ft 3 /h H 2 . 

Accordingly, dilution requirements can be calculated by: 

(5) (Ah/100) (number of cells) (0.016) 



Q(ff7min) = 



(0.008) (60) 



(5) (Ah/100) (number of cells) (0.016) 
60 



(15.4) 



where Q(ft 3 /min) is the minimum quantity of air necessary 
to dilute the H 2 produced to 0.8%, not allowing for dilution 
by room volume. To approach a worst case, it could be 
assumed that all batteries are 120-cell, 700 Ah, which is 
the present upper limit in battery size. Thus, each battery 
in tbe charging station requires about 140 ft 3 /min of air 
for good ventilation. 

This air quantity assumes that the battery lids re- 
main open or removed during charging. Although this is 
mandatory, it is a practice that is sometimes neglected, 
especially in low coal. For charging, vehicle storage bat- 
teries are placed in strong enclosures (trays) with remov- 
able lids. If these lids remain closed during charging, 
dangerous accumulations of hydrogen are likely to accu- 
mulate because of the small space above the battery top. A 
charging battery can be expected to liberate 60 ft 3 /h per 
25,000 cell ampere-hours. The result is an explosive con- 
centration unless forced ventilation is applied, and this is 
considered impractical by many. Enlarging the venting 
slots significantly is not an answer to the problem since it 
would weaken the battery box structure. The only solution 
is to ensure that the lids are open or removed. 



374 



BATTERY BOX VENTILATION 

The ventilation of the traction battery enclosure is not 
only important while charging but also during operation. 
Following the charge cycle, the covers are closed and the 
vehicle is placed in service. From chemical reactions and 
entrapped gas within the cells, lead-acid batteries con- 
tinue to emit gas for several hours after receiving a 
charge, be they open-circuited or on discharge (22). Con- 
sidering the ignitability of hydrogen, it is possible that a 
dangerous air mixture will accumulate. Hence it is neces- 
sary to provide adequate ventilation for the evolved hydro- 
gen in the closed box while the battery is in operation. 

Prior to 1945, there was little mention in the litera- 
ture of possible gas emissions from lead-acid batteries 
during discharge. In fact, many early publications stated 
categorically that there were no emissions under normal 
conditions (33), and this statement was repeated in mining 
literature (15). However, the possibility of gas emission did 
receive much attention from authorities responsible for 
high-capacity battery installations in such confined spaces 
as submarines. Robinson (22) quoted a report revealing 
that lead-acid batteries in confined spaces are always 
liable to emit considerable quantities of hydrogen and 
oxygen for the first few hours after charge. Although the 
actual quantities varied greatly from battery to battery, 
an 80-Ah cell, standing idle at 80°F (26.7°C) with 1.26 
specific-gravity electrolyte, would probably emit 5 to 20 
mL/h H 2 for about 12 h after charge. The approximate rate 
was found to be 

• Directly proportional to the battery capacity, 

• Doubled for each 15°F (9.5°C) rise in temperature, 
and 

• Doubled for each 0.050 unit increase in electrolyte 
density. 

As a result of antimony contamination of the negative 
plates, the open-circuit emission also increased with cell 
age. Upon discharge, additional hydrogen was evolved 
from the negative plates, which was thought to be a 
release of gas entrapped during charging. Combining the 
open-circuit and discharge emissions on a worst case basis, 
it was postulated that a lead-acid battery cell could release 
up to 500 mL (about 0.2 ft 3 ) of hydrogen per hour at the 
end of its useful life. 

An advisory committee on coal mining, formed in the 
United Kingdom during the mid-1940's, investigated such 
problems related to underground battery vehicles (28). The 
committee recommended that traction battery enclosures 
be properly ventilated to prevent a hazardous accumula- 
tion, and accordingly, a regulation for UK. coal mines was 
promulgated in 1949 (27). In the United States, the 
Bureau of Mines subsequently incorporated a similar 
requirement in Schedule 2G, stating that "battery boxes 
shall be adequately ventilated" (31). 

Corresponding to the 1949 regulations, a test proce- 
dure was made statutory for the United Kingdom (22, 29) 
in which 

• Hydrogen generation was calculated at 3.0 ft 3 /h per 
25,000 cell ampere-hours; 

• Tests were to be made in still air; 

• Maximum hydrogen concentration within the box 
under these conditions could not exceed 2.0% with tolerance. 



The emission calculations assumed that all factors con- 
tributing to the hydrogen-emission rate, such as cell 
temperature, acid specific gravity, and discharge rate, 
were simultaneously at worst case. Still air was selected to 
simulate a vehicle traveling at the same velocity as the 
mine ventilation so that the container's natural ventila- 
tion would receive no assistance. The maximum specified 
concentration provided a safety factor of 2.0 below the 
lower flammable level of hydrogen (4.0%). 

In his research, Robinson examined simple arrange- 
ments of container vents that would provide sufficient 
natural ventilation to meet these ventilation require- 
ments (22). He used the fact that hydrogen has the highest 
coefficient of diffusion into air of any gas; because of its 
extreme mobility, hydrogen is very difficult to retain 
within a leaky enclosure. The investigation used a simu- 
lated battery box containing a dummy 56-cell, 288-Ah 
lead-acid traction battery. Hydrogen liberation for dis- 
charge was calculated at 2.0 ft 3 /h by the foregoing rela- 
tionship. While hydrogen was pumped in at this steady 
rate, hydrogen concentrations were measured in the space 
between the battery top and cover (volume of 6.1 ft 3 ). 
Eleven different venting arrangements in the enclosure 
top were used, ranging from 40 to 140 in 2 , and sealed top 
vents were also investigated. End-plate venting was avail- 
able in all tests. 

Robinson's findings were as follows: 

1. Equilibrium between the battery rate of hydrogen 
emission and rate of hydrogen escape through the vents 
was established within 1 h after each test commenced. 
Thereafter, the hydrogen concentration remained almost 
constant. 

2. Maximum concentrations with natural top venting 
ranged from 1.3% to 2.8%, related inversely to vent area 
and following a near logarithmic curve. 

3. When the top vents were sealed and the side vents 
open, hydrogen concentration rose sharply to 5.3%, achiev- 
ing 4.2% in 8.0 min. However, with 31 ft 3 /min of forced 
ventilation through the side vents, the 5.3% maximum 
dropped to 1.7% H 2 . It was suggested that normal haulage 
speeds and mine ventilation flow rates would create a high 
factor of safety. 

The overall conclusion was that a battery box could be 
readily vented to meet U.K. requirements and forced 
ventilation was not needed. 

In 1959, Titman (26) reported on gas emissions from 
lead-acid batteries for the 45-min period after charging, 
supplementing his earlier study for alkaline cells. It was 
then suspected that gas emission rates immediately after 
charge might be greater than those emitted by the battery 
after a standing period. Titman also investigated the 
parameters causing variations in emissions, such as acid 
strength, discharge rate, and increased cell temperature. 
The experiments employed a new 6-V, 309-Ah traction 
battery. The conclusions of the research were 

1. Total gas emission for similar cells varied as much 
as 15%. 

2. Hydrogen emission rate doubled for an increase of 
0.050 in acid specific gravity (1.260 to 1.360). (Note: this 
increase could easily occur if the battery electrolyte was 
not topped up (19).) 

3. The rate doubled for a 12.5°C increase in electro- 
lyte temperature. 



375 



4. Overall rates were similar to those for alkaline 
cells. 

5. Immediately after charging, hydrogen rates up to 
5.0 L/h per cell were observed (corresponding to 14.30 ft 3 /h 
per 25,000 cell ampere-hours). 

6. After 45 min, the cells were found to liberate as 
much as 1.3 L/h per cell (3.72 ft 3 /h per 25,000 cell 
ampere-hours). 

7. The minimum emission rate was always reached 
within 5.0 to 8.0 min with the battery standing after 
charging. 

As can be seen, the findings for specific gravity were 
similar to those reported by Robinson. However, hydrogen 
emission rates were considerably higher than the 3.0 ft 3 /h 
testing standard: 376% higher immediately after charging 
and 24% higher after a brief standing period. 

Using these emissions rates as a basis, Titman (25) 
further investigated the effect of high hydrogen emission 
rates on typical battery box ventilation. The enclosures 
investigated were basically the same as those used by 
Robinson (22). The findings included 

1. A 2.0% H 2 concentration was not exceeded until the 
emission rate was 10 times the 3.0-ft 3 /h standard rate. 

2. After x /i h, the hydrogen became uniformly distrib- 
uted, with no tendency to accumulate at a specific place. 

3. For hydrogen concentrations of 2.0% or less, the 
enclosure concentration varied as the two-thirds power of 
the emission rate. (This was also verified theoretically.) 

4. With the high emission rates, acid-drainage holes 
assisted enclosure ventilation. 

This research has been presented in detail to demon- 
strate the reasoning behind battery box venting require- 
ments and to emphasize its importance. A direct compar- 
ison of hydrogen emission rates versus the top venting 
area of the enclosure would be advantageous, yet no direct 
relationship has been found. In practice, however, it has 
been discovered that about 25 in 2 of top vent area will 
allow 1.0 ft 3 H 2 to escape per hour, while meeting the U.K. 
testing requirements (20). As an alternative, the use of 
catalyst battery caps appears to be an excellent solution 
for difficult venting situations. 

These battery caps are a fairly recent introduction 
and are not yet used widely in mining, although they have 
been in use for some time in equipment ranging from 
torpedo batteries to personnel carriers in salt mines. They 
have also had some limited use in metal-nonmetal mining, 
such as uranium mines. In all these applications they have 
proved to be reliable and effective. 

A catalyst battery cap converts any emitted hydrogen 
and oxygen back into water (23). The caps have the dual 
function of preventing the escape of any hydrogen from the 
cells and restraining the loss of cell water. The technique 
has considerable safety advantages. No additional venti- 
lation is required for the tray or for the charging station 
itself. Battery lids do not have to be opened for charging, 
which is a major advantage in low coal. The explosion 
hazard associated with batteries is practically eliminated. 
Since watering of the cells is not required, no electrolyte 
can be spilled on the battery top, and the possibility of 
surface leakage (see below) is minimized. Tray corrosion 
from spilled electrolyte is similarly reduced. One problem 
with catalyst battery caps, however, is that the palladium 
catalyst may be destroyed if the cap is turned over. Special 
care must be taken not to do this during maintenance. 



BATTERY SURFACE LEAKAGE AND FAULTS 

A battery is designed to be an electrically floating 
system, insulated from its tray, which is at machine frame 
or ground potential. Yet three situations can occur that 
could connect the battery with its tray. 

1. Current may leak across the battery surface to the 
steel tray; this is referred to as surface leakage. 

2. A poorly insulated or damaged cable, bushing, and 
so forth may contact the tray, causing a fault condition. 

3. A rock fall or collision may force the battery box 
cover down onto the battery terminals, possibly shorting 
one or more to ground. 

Since lead-acid traction batteries are quite heavy, 
often weighing several hundred pounds, the only material 
presently available for tray construction is heavy-gauge 
steel. Steel is a reasonable conductor of electricity, a 
disadvantage when it is used to support an isolated system 
such as a battery. After a battery has been in operation for 
some time, dust mixed with spilled electrolyte can collect 
on the battery top. If these contaminants are permitted to 
accumulate, a number of low-resistance paths may form 
between the various cell terminals and the tray. Since the 
terminals are at different potentials, currents tend to 
circulate across the battery top and through the tray, and 
may cause the following problems: 

1. Currents circulating in resistive loops cause heat- 
ing. This heating is proportional to I 2 R, where I is the 
leakage current and R is resistance of leakage path. When 
R is low enough to permit a substantial current to flow, a 
smoldering fire may ensue. This hazard is compounded by 
the often explosive hydrogen concentrations that occur in 
the battery above the electrolyte level. 

2. The presence of paths from the cell terminals to the 
tray can be a shock hazard for mining personnel. A miner 
leaning against a battery box and touching an exposed 
terminal could be shocked. Terminal-to-tray conducting 
paths present an especially serious hazard when the 
battery is charging if the charger design or a fault within 
the charger permits dc to flow in the frame or ground. 

3. Circulating currents waste battery power and may 
reduce the amount of time a battery can be used before 
recharging is necessary. They also tend to increase tray 
corrosion. 

Another potentially hazardous situation can be 
caused by a low-resistance ground fault, for example, 
produced by a damaged cable in contact with the battery 
tray. Since the battery is theoretically a floating system, a 
single such fault should cause no current flow, but the 
effect of surface leakage paths can result in a certain level 
of current, depending upon path resistance. Two simulta- 
neous low-resistance ground faults, however, would permit 
extremely high currents to flow. Depending on the contact 
resistance of the fault, currents on the order of 10,000 A 
may exist for a short duration because of the low internal 
resistance of lead-acid batteries. A very hazardous situa- 
tion could result because batteries cannot be deenergized; 
hence, a fire caused by faults would be difficult to extin- 
guish until the battery was discharged. 

The surface leakage problem has been recognized by 
battery manufacturers for some time, and they constantly 
expound the virtues of keeping battery tops clean. An 



376 



example is the following excerpt taken from a typical 
maintenance article (11): 

If the battery tops become wet and dirty, or if tray 
corrosion is visible, give the battery a soda wash. 
Mix a handful of bicarbonate of soda (baking soda) 
in a bucket of water. Pour this solution over the top 
of the battery, using one full bucket per tray. Be 
sure vent caps are in place. Water will dry, leaving 
some dry soda on the battery top. It is good practice 
to give batteries a soda wash once a month. If a 
battery is accidentally flooded (acid spilled on cell 
tops) due to overfilling cells, give soda wash as soon 
as possible. 

Indeed, if mine batteries were always kept clean and dry, 
the surface leakage problems would be greatly reduced. 
Unfortunately, cleaning is often neglected in the mining 
industry, especially in low coal where batteries are so 
difficult to access. It would therefore be advisable to find 
some alternative way of reducing this problem. 

Most batteries now in use underground have a thick 
coating of paint on their trays. This reduces leakage 
somewhat, but the paint is prone to deterioration by 
chipping, abrasion, and attack by battery acid. Recently, 
some manufacturers have been coating their battery boxes 
inside and out with a tough vinyl compound sold under the 
brand name Plastisol. This product is readily available 
and can be sprayed onto any properly prepared steel 
surface. It appears to reduce surface-leakage problems 
significantly. 

Battery safety could be increased further if the ex- 
posed intercell connectors were insulated in some manner. 
Unfortunately, this is not a straightforward matter since 
any insulating coating applied to the connectors reduces 
their heat-transfer capabilities and the intercell connec- 
tors remove about 80% of the battery heat. 

As an option on their mining batteries, leading bat- 
tery manufacturers now produce a dead-top battery with 
totally insulated intercell connectors. This system should 
greatly reduce the potential for surface leakage and also 
remove the possibility of dangerous arcing if a tool is 
dropped across the battery top. 

Since surface leakage and ground faults are potential 
hazards, a method is needed to detect and isolate them. 
Recognizing this, Statham and Littlewood (24) developed a 
fault-detection system designed to be fitted to the battery 
or battery charger. This system, shown in the circuit in 
figure 15.12, is a simple way to detect ground faults that 
might occur between the battery and its load. One problem 
with the circuit is that a large portion of the battery itself 
cannot be protected: for instance, a fault occurring at point 
A in figure 15.12 would produce no current flow through 
the current relay. Figure 15.13 is a plot of relay current 
versus fault position on the battery; it is assumed that the 
current-relay shunt is set at 1,000 fi and that a 200- V 
traction battery is used. In order to increase the sensitiv- 
ity of their circuit to battery surface leakage faults, 
Statham and Littlewood proposed installing switches be- 
tween point X and frame and between point Y and frame 
(fig. 15.12). These switches would be alternately opened 
and closed by a mechanical or solid-state device. The upper 
dashed lines in figure 15.13 show the new position- 
sensitivity curve resulting from the modified circuit. The 
entire battery could be protected in this manner, although 
sensitivity still varies with fault position for surface 
leakage faults. 




Figure 15.12.— Circuit for detecting faults in batteries. 





70- 


>^ Relay characteristics 


9 


r -70 




60- 


>. with switches between points , 


♦ 


- -60 


<-( 




v. X and frame or ,♦* 






E 


50- 


>v Y and frame ,♦* 




- -50 


h- 










7* 




^^ ♦ 






Ld 


40- 


\^ ♦ 




- -40 


or 




^V ,* 






or 

=) 
o 




^v ♦* 






30- 


N. y >. 




- -30 


V 




^V. * ^^. 






< 




^V ♦* ^« f 






_1 


20 - 


^V. * ^V. m 




20 


Ld 

or 




\^* Original circuit yf 

♦* >^ characteristics^ \^ 








10- 






- -10 




C 


I T i 

1 50 100 150 


200 



ASSUMED FAULT POSITION 
RELATIVE TO BATTERY VOLTAGE, V 

Figure 15.13.— Curve of relay current for various fault posi- 
tions on battery. 

Virr and Pearson (34) devised an electronic ac injec- 
tion system that would provide protection sensitivity in- 
dependent of fault location. They utilized a system of red 
and green lamps to indicate a not safe or safe condition. 
Although Virr and Pearson recommended mounting the 
device on each battery, this would not be necessary for 
trackless battery-powered vehicles since a single device on 
the charger would be sufficient to prevent an unsafe 
battery from being charged. Since the development of 
surface leakage presumably takes place gradually with 
the accumulation of conducting material on the battery 
top, continuous monitoring of each individual battery is 
probably unnecessary. 

There is an additional hazard when using lead-acid 
batteries in any environment: an explosive hydrogen-air 
mixture can be available inside the cell. This is true even 



377 



with catalytic caps. If ignition energy is sufficient within 
the cells or even close to the vent cap, the internal mixture 
can explode, possibly spraying acid and blowing bits of the 
cover toward personnel in the vicinity. Internal faults can 
produce this kind of explosion. Virr and Pearson related 
that a common source of such events also occurs when 
electrolyte leaks from a cracked cell, producing a spark 
when the acid level falls below the bottom of the plates. 
There is no known method of preventing or suppressing 
these internally initiated cell explosions. However, 
ground-fault protection will detect electrolyte leakage. 

Low-resistance faults to the tray or vehicle frame 
caused by damage to cables, and so on, represent a 
different problem. As previously mentioned, no provision 
is currently available for deenergizing a faulted battery, 
which would continue to discharge until its stored energy 
was dissipated. Some form of circuit breaker between the 
cells of the battery would provide a way to sectionalize the 
battery in the event of a fault. In conditions of excessive 
current flow, the intercell connectors might themselves act 
as protective fuses, melting down when their current- 
carrying capacity was exceeded. 

Proper isolation of the battery electrical system from 
the mine ground system is a prerequisite for safe battery 
use. The reliability of battery isolation can be greatly 
enhanced by following the maintenance and design sug- 
gestions given in this chapter. Total electrical safety for 
any battery installation is, however, a function of charger 
characteristics as well as battery isolation. 



BATTERY-CHARGING HAZARDS 

A brief review of four representative accidents sheds 
some light on the hazards associated with battery 
charging: 2 

1. A scoop operator was electrocuted when his body 
came in contact with the frame of a scoop tractor that was 
being charged. During the subsequent investigation, it 
was found that a potential difference of 260 V existed 
between the tractor frame and mine floor when the 
charger was energized. The cause was a low-resistance 
surface-leakage fault between the battery and battery 
tray, which caused the tray and the tractor frame on which 
it was resting to become energized. Faulty insulation 
between the primary and secondary windings on one arm 
of the three-phase transformer permitted secondary cur- 
rent to flow in the ground. The glaring error here was the 
failure of mine personnel to ground the steel frame of the 
scoop tractor properly while it was being charged. 

2. A utility worker received a fatal electric shock 
when his body contacted the frame of a battery charger. A 
fault occurred within the charger that caused 210 Vdc and 
114 Vac to exist between the charger frame and earth. 
Despite the fact that the primary cause of the electrocu- 
tion was a worn bushing that failed to insulate the timer 
circuit from the charger frame, a proper frame ground 
could again have prevented the fatality. 

3. An electrician was fatally injured when he came in 
contact with a bare conductor on the charging leads of a 
battery charger while connecting the charger to the vehi- 
cle. The electrician was standing in water while attempt- 
ing to connect the battery. Although the charger switch 



2 MSHA reports of fatal coal mine electrical accidents. 



was in the off position, a primary-to-secondary fault in the 
charger transformer circumvented the switch (which in- 
terrupted only one primary conductor) and caused the 
charger leads to become energized. 

4. A surveyor was electrocuted when he contacted the 
battery ground clamp that was attached to the charger 
frame. The charger frame was energized because of a fault 
within the charger. The investigation showed the accident 
to be due to inadequate safety grounds on the frames of all 
the electrical equipment in the mine. 

These accidents occurred using chargers containing 
the basic circuitry discussed earlier in this chapter, and 
with the possible exception of the enclosure, they were 
similar to chargers used by other industries. This implies 
that additional design concepts or components must be 
included in battery chargers in order to ensure safety in 
the mine environment. 

Poor grounding and component failure (most espe- 
cially the power transformer) have been the most notori- 
ous contributors to mine charger accidents and electrocu- 
tions. They have caused the ac source power to be 
impressed on the dc charging circuit. Consider an instance 
where a primary-to-secondary power transformer fault 
elevates the charging circuit by the primary potential. If 
this occurs simultaneously with a fault that energizes a 
battery tray, and the grounding is unsatisfactory, the 
battery tray will be a shock hazard. Of vital importance is 
adequate grounding because an intact grounding system 
ensures that frame potentials do not exceed reasonably 
safe levels, regardless of the contribution of other factors to 
the hazardous condition. 

A list of other problems causing accidents has been 
made through an analysis of battery charger accidents and 
input from operators, manufacturers, and State and Fed- 
eral regulatory agencies: 

• Bad charging-cable insulation; 

• Unconnected, exposed charging couplers (plugs) 
that are still energized; 

• Poorly designed charging couplers that have ex- 
posed ungrounded metal parts where the cables are at- 
tached and allow cable damage; 

• Charging couplers with poor contact as a result of 
frequent damage or inadequate maintenance; 

• Battery surface leakage and internal faults; 

• No overcurrent protection for the ac power input; 

• Faults causing the control circuitry and compo- 
nents to have an elevated potential; 

• Personnel connecting a charger to a battery of 
wrong polarity; 

• Repair work performed by unauthorized personnel; 
and 

• Charging stations in abnormally wet locations. 

In view of this poor safety record, it is advisable to 
consider the safety features that would reduce the potential 
for accidents and electrocutions when charging batteries: 

1. The charger input cable should contain a monitored 
ground to ensure that the grounding conductors to the 
charger frame are intact. 

2. The charger should be equipped with panel inter- 
locks that deenergize the charger at the outby source when 
access panels are removed. 



378 



3. The charger should have an emergency off switch 
located in a conspicuous place on the charger frame. This 
switch should not be spring-loaded, thus requiring reset- 
ting after use. 

4. The power transformer should electrically isolate 
the battery being charged from the power source, with 
primary and secondary windings being so arranged to 
eliminate hazardous interwinding faults. 

5. A separate grounding conductor and ground-check 
monitor circuit should be provided for each battery tray 
serviced by the charger. 

6. The dc couplers should be of the type that inter- 
rupts the ground-check circuit before the charging circuit. 

7. The dc connection between the charger and battery 
box should consist of a single cable with appropriate 
grounding and ground-check conductors. 

8. The charger should contain battery surface leakage 
detection circuitry that prevents a leaky battery from 
being charged. 

9. The power transformer secondary and all dc circuits 
and components should be isolated from the frame ground 
of the charger. 

10. The charger should have a meter or similar device 
that indicates the state of battery charge. 

11. The charger should have overload and short- 
circuit protection on both the input and output. 

12. The charger should contain circuitry that prevents 
a battery of the wrong voltage from being charged; alter- 
natively, the charging connector must be keyed or sized 
such that no connection can be made with a battery with 
fewer cells than that for which the charger is designed. 

Figure 15.14 is a diagram of a typical solid-state 
controlled taper-rate charger. All the desired electrical 
components specified in the above list are signified by the 



letters in parentheses. These and other features will be 
discussed in the following paragraphs. It should be noted 
that in some instances there are alternative practical 
methods that could be employed to protect personnel. 

System Grounding 

Because of its ac-to-dc application, the charger illus- 
trated is considered to be portable electrical equipment 
and is fed by a resistance-grounded low-voltage or 
medium-voltage ac system. Under Federal regulations for 
coal mining, a grounding conductor and ground-check 
monitoring of this grounding conductor are required be- 
tween the power source (usually a power center) and the 
charger frame. The portion of the ground-check circuit 
within the charger is indicated at a. The panel interlock 
switches (o) and the emergency off switch (c) are shown in 
series with the ground-check circuit. Here, a pilot-type 
monitor is implied. If the circuit is opened in any manner, 
the circuit breaker at the power source is tripped. After 
tripping, the circuit breaker must be reset manually at the 
source and cannot be reset until the ground-check circuit 
is restored. 

Electrocutions involving contact with the charger 
frame have taken place when grounding practices were 
unsatisfactory and a fault existed between the frame and 
some charger internal component. One method for pre- 
venting this type of hazard would be to use an insulation 
coating to isolate the enclosure completely from possible 
contact with live circuits. There are, however, several 
shortcomings to this method of protection. Most battery 
chargers designed for underground use are mounted on 
skids for easy mobility. In rough mine use, any insulation 
would be quickly worn off, especially on the skids. Insula- 
tion coatings would be difficult to apply everywhere. 



Incoming 
cable 




check 
Ground 



<<e 



[a) 



Fuse 



n 



kg) 

-= Faraday 
shield 



[b) 

Panel 
interlock 
switches 

(c) 

Emergency 

off 




Control 
transformer 



Ground-check 
monitor 



(/) 

dc 
contactor 

HI — 

K3 

HI — 



Control 
circuits 



X 



(/?)=j:K1 

(/)^K2 
/ ;\ -L Manual 
U> Toff -on 

(<?)(m j Main timer 



Cable to 

battery 

tray 

-» — Positive 



-» — Negative 




Ground 
check 



Ground 



Figure 15.14.— One-line diagram of desired charger features. 



379 



Devices mounted to the cabinet would have to be coated for 
100% protection, and bonding all parts together and 
bonding components like timers and switches to the ex- 
posed panels would pose a considerable problem. Further- 
more, any insulating compound deteriorates with time. 
Thus insulation coating(s) might both be impractical and 
result in a false sense of security. As a result, present 
enclosure grounding standards may well be a more effec- 
tive safety approach. 

Panel Interlocks 

The main reason behind the panel interlock switches 
is to prevent authorized or unauthorized personnel from 
unknowingly contacting live parts inside the charger. The 
emergency stop or "panic button" provides definite safety 
advantages. In case of an emergency, it provides a quick 
power stoppage at the charger. Without such a switch in 
an underground mine, a miner would have to go back to 
the power center, which could be as far as 500 ft away. The 
switch has a large red button for easy identification; 
pushing the button breaks the upstream ground monitor 
circuit. The button must be pulled out manually before the 
outby circuit breaker can be reset and so provides a double 
check on reenergization. This is particularly important for 
maintenance personnel since it prevents power from being 
accidentally restored to a circuit while they are working 
on it. 

A justifiable complaint against panel interlocks is 
that they must be defeated for maintenance and may be 
left in that condition, rendering the system dangerous, 
from a false sense of security. However, experience with 
panel interlocks on other mine power equipment has 
shown that these safety devices rarely remain in a de- 
feated condition. Remember that any protection system 
can be knowingly or unknowingly rendered useless, and it 
is the responsibility of training personnel to make sure 
this is minimized. There are other ways to discourage 
unauthorized entry; for example, barriers or partitions 
can separate energized components within the charger, 
and special opening tools and attachments can be used on 
many covers. But even in these cases, panel interlocks 
would give an extra safety margin for any panel that is 
opened or removed for normal adjustments. 

Transformer Failures 

A Faraday shield is indicated at g in figure 15.14. This 
prevents primary-to-secondary (interwinding) faults in the 
power transformer and has the secondary benefit of isolat- 
ing any high-frequency transients on the incoming power 
from the charger where they could interfere with the 
control circuitry. If the amplitude of these transients was 
high enough, they could destroy solid-state devices. The 
Faraday shield probably exceeds the desired safety re- 
quirements for secondary-circuit isolation. 

Axial displacement of the secondary and primary 
windings on the transformer core plus separation of pri- 
mary and secondary circuits is also acceptable instead of 
the grounded shield. Again, the goal is to accomplish 
electrical isolation of primary and battery circuits. 

Transient overvoltage conditions on mine power sys- 
tems, caused by switching or in some cases lightning, 
apply large electrical stresses to power transformer insu- 
lation. To protect the transformer and also to provide 
backup for the personnel protection given by the Faraday 



shield, the transformer insulation should be able to with- 
stand a peak of five times the nominal line voltage peak. 
Ventilated dry-type transformers designed to IEEE stan- 
dard 462-1973 (16) for low-voltage chargers have a 10-kV 
BIL, which is well within the transformer transient- 
protection needs. Accordingly, the voltage ratings of all 
other charger components are coordinated with the maxi- 
mum anticipated overvoltage. The transformer secondary, 
and thus all circuits connected to it (rectifiers, charge-rate 
control, and timer circuitry) are further protected from 
transients by surge traps (metal oxide varistors) fin figure 
15.14. 

Outgoing Cables 

The cable used to charge the batteries contains 
grounding and ground-check conductors as well as the two 
charging power conductors. The battery tray couplers 
should have four pins to accommodate the cable. Obvi- 
ously, the general safety requirements for any low- voltage 
coupler should also be adhered to: for example, each 
contact should be able to continuously carry the maximum 
current in the circuit for which it is designed; all exposed 
uninsulated metallic parts should be grounded to the 
grounding conductor; and the coupler-cable interface 
should be designed to prevent cable insulation damage. 

The grounding and ground-check conductors are 
grounded on separate welded studs inside the battery tray. 
If only one stud is used and this is knocked loose, the 
ground-check circuit could be intact but the battery tray 
could be ungrounded. 

A logical safety item is provision of insulated strain 
relief for all cables associated with a charger. At least one 
previous fatality could have been prevented if cable strain 
relief had been provided. To increase protection against 
cable damage, insulated glands should be provided for any 
cable passing through the charger frame, and there should 
be adequate storage on the outside of the enclosure for all 
permanently attached cables. This would alleviate dam- 
age caused by leaving charging cables and couplers on the 
mine floor. 

Outgoing Ground-Check Monitoring 

The ground-check circuit designated by h in the figure 
monitors the grounding connections to the battery tray: 
one circuit is provided for each battery on charge. The 
monitor shown is a simple current-sensing pilot or loop 
monitor. If the loop formed by the charger cable grounding 
and ground-check conductors, the coupler contacts, and 
the battery tray grounding studs is opened, relay Kl will 
be deenergized and trip the charging power with contactor 
K3. A serious problem with ground-check monitoring on 
battery charging systems is the potential for development 
of parallel grounding paths. For this and other reasons, 
the monitor must be designed to meet all the guidelines 
discussed in chapter 9. 

Surface Leakage Detection 

The charging power will also be tripped if the leakage 
detector circuit (i) senses surface leakage or a ground fault 
from the battery to the tray. Following work performed by 
Virr and Pearson (34) in the United Kingdom, a total 
battery leakage resistance threshold of 1,000 fi appears 
satisfactory. This value should not have nuisance tripping 



380 



problems caused by too-sensitive relaying, but should 
substantially reduce the gas ignition and electrical shock 
hazards created by such faulting. 

Charging Circuit Isolation 

The goal of completely isolating the transformer sec- 
ondary and dc charging circuitry from ground is aimed 
mainly at preventing shock hazards caused by charger- 
battery grounding problems. With either polarity 
grounded, a low-impedance source of dc voltage is avail- 
able between the frame of the vehicle being charged and 
the ungrounded battery side. A person touching the vehi- 
cle frame and any battery intercell connector could then 
receive a dc shock. The same would be true for anyone 
standing on a wet mine floor and touching any un- 
grounded terminal connected to the charging circuitry. If 
the positive or negative charger terminal is grounded, this 
would also promote excessive corrosion of the battery tray. 

State-of-Charge Indication 

An ammeter (k) is provided so that at any given time 
it is obvious what part of the charge cycle the system is 
operating in. Strictly speaking, this is not a safety feature, 
but having knowledge of the status of the charge cycle is 
an aid to good operating procedures, which in turn can 
contribute to safety. 

Another valuable feature would be a fail-safe means of 
preventing overcharge. The principal goal here is to pro- 
tect against excessive battery gassing, and as a first line of 
defense, almost all manufacturers use timing circuitry to 
reduce the charging current substantially through the last 
part of the charge cycle. However, there are common 
occurrences that call for additional protection. For in- 
stance, consider that the charge cycle is interrupted dur- 
ing charging, say by a charger component failure, a power 
failure, or someone disconnecting the charger from the 
batteries and then reconnecting them. The bad component 
could cause the charger to continue the charge at the 
starting rate for an indefinite period. In the last two cases, 
the main timing circuitry could reset to zero, and an 
unknowing individual might manually restart the 
charger for a full charge cycle. Considering the real world, 
a "fail-safe" timer would need to deenergize the charger 
with the occurrence of any failure mode that could lead to 
overcharge. Many of these failures are perceivably beyond 
the control of practical timing circuits; thus, a more 
reasonable safety requirement would be having the 
charger time circuitry operate to minimize battery over- 
charge as much as practical. 

One means of affording protection against overcharg- 
ing would be to have a redundant timing device that would 
override the main timing circuitry and shut the charger 
down after a set period. The timers (e) in figure 15.14 
perform this function: their dc motors are tied to the 
charging power and have automatic reset capability. The 
time deenergizes the charger after a maximum opera- 
tional period measured from its initial setting. 

Overcurrent Protection 

Fuses id) and it), in series with each ungrounded 
incoming and outgoing power conductor, provide overload 
and short-circuit protection. A circuit breaker could also 
be used for the transformer primary circuit, but this is 



probably not suitable for outgoing-circuit protection. 
Other relaying could also be included for semiconductor 
protection, as discussed in chapters 12 and 14. 

Additional Features 

The preceding discussion covered the minimum num- 
ber of features that should be included in a mine charger, 
but some additional items are also desirable. One of these 
is a contactor (i) for both the positive and negative outgo- 
ing conductors, which should be located near the point 
where charging power leaves the enclosure. This would 
remove battery power from any uninsulated component 
within the enclosure if the charger is opened and the 
charging cable is still connected. Accordingly, the contac- 
tor would trip any time the ac ground-check circuit is 
broken. Because the contactor load contacts might still be 
energized upon entry, the device should have an insulated 
cover with a warning notice. 

There have been many instances where charging 
couplers have been removed during the charge cycle, with 
the charger not turned off. The energized plug has been 
left on the mine floor, at times lying in water. The outgoing 
ground-check circuit will prevent this, as well as any 
arcing that might be encountered during plug removal. As 
a backup to power tripping, it would also be advantageous 
to have the manual on-off device simultaneously reset to 
off. This could be a simple switch tripped by the dc 
ground-check monitor or a manual timer that is capable of 
being automatically reset to zero (component j in figure 
15.14). Automatic restarting of the charger (for example, 
when the battery plug and receptacle are reengaged) is 
inadvisable as it could pose a hazard to the unwary miner. 
Manual restarting is preferred as this necessitates a 
deliberate act by the user before the charger can be 
energized. 

A very serious and quite obvious hazard can also 
occur if an energized charger is connected to a battery of 
wrong polarity. It is essential that the control circuitry be 
able to sense incorrect polarity and if it exists, prevent 
energization of charging power. Many maintenance per- 
sonnel have received burns from battery-energized cou- 
plers during repair, maintenance, or replacement. Hence 
it is necessary that the battery terminal connections be 
constructed so power to the couplers can be removed. 

This chapter has scanned the subject of batteries and 
battery charging in mining. Some significant points have 
been covered. A good battery-maintenance program is 
essential for successful battery use. Adequate ventilation 
of the typical charging station requires little more than 
common sense. The batteries must be located in the 
mainstream of airflow during charging, and provisions 
should be made to deliver 140 ft 3 /min for each battery on 
charge. Battery lids should be removed during charging. 
Catalyst battery caps might be considered as a viable 
alternative to ventilation and lid removal for handling 
hydrogen accumulation problems. Safe battery chargers 
can be made and should he used to provide a larger safety 
factor in the battery charging and usage process. 

The Federal requirement to deenergize the mine elec- 
trical power system in the event of a mine ventilation 
failure raises an unresolved question concerning battery 
employment. System deenergization should include all 
nonpermissible battery-powered systems at the source. 
Because batteries are independent and cannot sense a 
mandatory power shutdown, a "dead-man" type of device 
may be needed on battery-powered vehicles. This device 



381 



would be timed to interrupt the battery power, perhaps 20 
min after the machine is stopped. The operator would be 
required to engage an interlock before the machine could 
be reactivated. Such a device could be interlocked through 
the battery tray coupler, thus deenergizing the receptacle 
contacts when the plug is not present. 

REFERENCES 

1. Appleton, J. Storage-Battery Locomotives in Mine Work. 
Coal Age, v. 16, Nov. 16, 1919. 

2. Burkle, B. J. The Use of Traction Type Storage Batteries 
Underground in British Mines. Min. Technol., v. 53, Jan. -Feb. 
1971. 

3. C & D Co., Batteries Div. (Plymouth Meeting, PA). General 
Service Manual -Motive Power Batteries. Undated. 

4. How To Get Longer Life From Motive Power Bat- 
teries. Coal Age, v. 80, Oct. 1975. 

5. Coal Age. Storage-Battery Locomotive Arc Said To Have 
Started Everettsville Explosion. V. 31, June 16, 1927. 

6. Exide Co., Power Systems Div. (Philadelphia, PA). The 
Power Maintenance, Repair, and Safety Procedures of Motive 
Power Batteries. 1974. 

7. Fink, D. G., and J. M. Carroll (eds.). Standard Handbook for 
Electrical Engineers. McGraw-Hill, 10th ed., 1968. 

8. General Battery Corp. (Reading, PA). Industrial Battery Ser- 
vice Manual. 1975. 

9. Gould, Inc., Industrial Battery Div. (Trenton, JN). Instruc- 
tion, Maintenance, and Service Manual, Motive Power Batteries 
and Chargers. 

10. Harvey, R. A. Battery Chargers and Charging. Iliffe and 
Sons, Ltd., London, 1953. 

11. Hensler, J. F. Battery Care for Mine Tractors. Coal Age, v. 
70, Nov. 1965. 

12. The Case for Batteries. Coal Age, v. 70, Mar. 1985. 

13. Hobart Brothers Co., Power Systems Div. (Troy, OH). Fer- 
roresonant Charger Transformer Theory. Data Sheet 2715, July 9, 
1970. 

14. Hopewell, R. W. Safety Check List for Batteries. Saf. 
Maint, v. 121, Feb. 1961. 

15. Ilsley, L. C. Development and Safety of the Storage-Battery 
Locomotive. BuMines IC 6068, 1928. 

16. Institute of Electrical and Electronics Engineers (New 
York). Recommended Practice for Electric Power Distribution for 
Industrial Plants. Stand. 141-1986. 

17. KW Battery Co. (Skokie, IL). Installation and Maintenance 
Manual -KW SCR Lifeguard Charger. May 1975. 



18. 



Service Manual for Motive Power Batteries. 1975. 



19. Morley, L. A., and J. A. Kiefer. Coal Mine Electrical System 
Evaluation (grant G0155003, PA State Univ.). Volume V-Battery 
and Battery-Charging Safety. BuMines OFR 61(5)-78, 1977; NTIS 
PB 283 494. 

20. Neill, A. G. The Lead-Acid Battery With Particular 
Reference to Its Use in Mining. Min. Electr. and Mech. Eng., v 4b, 
Nov. 1965. 

21. Owings, C. W. Hauling Coal Safely With Permissible 
Storage-Battery Locomotives. BuMines RI 3051, 1930. 

22. Robinson, H. The Ventilation of the Battery Containers of 
Storage Battery Locomotives. SMRE, Sheffield, England, Res. 
Rep. 122, Feb. 1956. 

23. Sanford, J. Plastic Caps Reconvert Battery Gases to Needed 
Water. Electron Des. News, v. 20, Nov. 3, 1975. 

24. Statham, C. D. J., and J. Littlewood. Developments in the 
Safe Use of Traction Batteries in Coal Mines. Paper in IEEE Con- 
ference Publication 74. 1971. 

25. Titman, H. The Effect of Different Rates of Emission on the 
Accumulation of Hydrogen in a Vented Battery Container. SMRE, 
Sheffield, England, Res. Rep. 167, Mar. 1959. 

26. The Emission of Gas From Lead- Acid Cells. SMRE, 

Sheffield, England, Res. Rep. 158, Feb. 1959. 

27. U.K. Ministry of Fuel and Power (now U.K. Dep. Energy). 
Coal Industry: The Coal Mines (Locomotives) General Regulations. 
Statutory Instrum. 530, H.M.S.O., London, 1949. 

28. Coal Mining: Report of the Technical Advisory Com- 
mittee. Command Paper 6610, H.M.S.O., London, 1945. 

29. Tests and Specifications for Storage Battery 

Locomotives for Use in Mines Under the Coal Mine Act, 1911. Test. 
Memo. 11, H.M.S.O., London, 1949. 

30. U.S. Bureau of Mines. Schedule 15, Procedure for 
Establishing a List of Permissible Storage-Battery Locomotives 
for Use in Gaseous Mines: Character of Tests, Conditions Under 
Which Storage-Battery Locomotives Will Be Tested and Fees. 
1919. 

31. Schedule 2G, Electric Motor-Driven Mine Equip- 
ment and Accessories. Federal Register, v. 33, No. 54, Mar. 19, 
1968. 

32. Vaughan, K. A. Storage Battery Charging and Modified 
Constant-Voltage Method. Mechanization, v. 15, June 1951. 

33. Vinal, G. W. Storage Batteries: A General Treatise on the 
Physics and Chemistry of Secondary Batteries. Chapman and Hall, 
Ltd., London, 1924. 

34. Virr, L. E., and F. K. Pearson. Fail-Safe Earth-Fault- 
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1974. 



382 



CHAPTER 16.— PERMISSIBILITY AND HAZARD REDUCTION 



Any industrial area in which flammable or explosive 
gases, vapors, and dust can be encountered is designated 
as a hazardous location. Since the occurrence of a hazard 
depends upon the presence of an ignitable mixture, an 
ignition source, and contact between them, the chance of 
ignition is always present when electrical apparatus is 
used in hazardous atmospheres. The probability of igni- 
tion cannot be brought to zero in any portion of an 
underground mine, nor in selected parts of surface mines 
and surface facilities such as preparation plants. Hence 
electrical hazard reduction techniques must be applied in 
these areas to protect both personnel and equipment. 



TERMINOLOGY 

Important measures used to reduce incendive hazards 
in underground mines include provision of adequate ven- 
tilation, control of flammable coal dust through manda- 
tory rock dusting and watering, and the regulation of 
equipment. The last item is the one of greatest interest 
here. The regulation of electrical face equipment in under- 
ground mines is specified in 30 CFR 18 (37). 1 

The responsibility for ascertaining compliance with 
this document is assumed by the U.S. Department of 
Labor, Mine Safety and Health Administration (MSHA), 
Approval and Certification Center. It is important that the 
terminology for Part 18, "Electric Face Equipment," is 
clearly understood. 

Approval: This term applies to completely assembled 
electrical machines and accessories. Accessories mean 
associated electrical equipment such as a distribution or 
splice box that is not an integral part of the machine. 
Approval means that a formal document has been issued 
by MSHA, which states that the machine or accessory has 
met the applicable requirements of the regulation. An 
approval plate is then attached to the approved machine or 
accessory identifying it as suitable for use in hazardous 
locations. Such equipment is subsequently referred to as 
permissible equipment. This process is mandatory for all 
electrical equipment used inby the last open crosscut and 
in return air of an underground coal mine. 

Certification: This term applies to an electrical com- 
ponent, that is, an integral part of an electrical machine or 
accessory that is essential to the functioning of the ma- 
chine. Certification means that a formal written notifica- 
tion has been issued by MSHA, which states that the 
component complies with Federal requirements and is 
suitable for incorporation in a permissible machine. 

Acceptance: This term applies to flame-resistance re- 
quirements and to auxiliary equipment such as a cable, 
hose, or belt. Acceptance means that written notification has 
been received from MSHA designating the equipment as 
meeting requirements for flame resistance. Acceptance 
marking is the identification that appears on the equipment. 

Intrinsically safe: This identifies equipment that is 
incapable of releasing enough electrical or thermal energy 
under normal or abnormal circumstances to cause ignition 
of a flammable mixture. 



1 Italicized numbers in parentheses refer to items in the list of references 
at the end of this chapter. 



Hazardous locations in surface mines and surface 
portions of underground mines are often classified by 
guidelines specified in the National Electrical Code 
(NEC), article 500 (23). In this classification system, the 
nature of the hazard and the degree of hazard are the main 
considerations, and a location is specified by a class, group, 
and division designation. The class refers to the generic 
nature of the hazardous material, and the following are of 
specific importance in mining: 

• Class I: Locations containing flammable gases or 
vapors that may be present in the air in sufficient quantity 
to produce an explosive or ignitable mixture. 

• Class II: Locations having combustible dust in 
quantities that can cause a hazard. 

The group designation is a subclassification and refers to 
the nature of the hazard. Each group contains a listing of 
materials that present the same general hazard. The 
groups of interest in coal mining could include 

• Group B: Atmospheres containing hydrogen or 
gases or vapors of equivalent hazard such as manufactured 
gas. 

• Group D: Atmospheres containing gasoline, hex- 
ane, naptha, benzine, butane, propane, alcohol, acetone, 
benzol, lacquer-solvent vapors or natural gas (methane). 

• Group F: Atmospheres containing carbon black, 
coal, or coke dust. 

When more than one hazard is involved, the class and group 
describing the most serious situation usually applies. 

The division defines the probability of a hazardous 
material being present in an ignitable concentration: 

• Division 1: Locations where the hazards exist con- 
tinuously, intermittently, periodically, or where they may 
exist during maintenance or equipment failures. 

• Division 2: Locations where hazards are presumed 
to exist only under abnormal conditions. 

An example of a division 2 location is an area rendered 
nonhazardous by forced-air ventilation that could become 
hazardous if the ventilation failed. Another instance 
would be a location where dust layers could accumulate 
and might interfere with proper and safe heat dissipation 
from electrical equipment, thereby causing a dust-layer 
ignition. 

The NEC, article 501, contains the general rules 
applicable to electrical wiring and equipment in class I 
locations, and article 502 applies to class II locations. 
Electrical equipment used in preparation plants com- 
monly follows design criteria for hazardous locations es- 
tablished by organizations other than MSHA, but MSHA 
has jurisdiction. The National Electrical Manufacturers 
Association (NEMA) defines criteria for explosion-proof 
and dust-ignition-proof enclosures, motor classifications, 
and insulation classifications (22). Underwriters' Labora- 
tories, Inc., lists motor and control standards for hazard- 
ous locations (35). 



383 



HAZARD-REDUCTION METHODS 

Hazard reduction is a practice in low probabilities, 
specifically low incremental probabilities, and relies on 
the fact that a safe electrical installation in a hazardous 
location does not significantly raise the probability of fire 
or explosion above that existing without the equipment 
(14). The methods adopted to reduce hazards are based on 
the principle that the occurrence of a hazard depends upon 
the presence of an ignitable substance, an ignition source, 
and contact between them. The methods focus on one or 
another of these interrelated factors. 

Among hazard-reduction methods used in or around 
mines, explosion-proof and dust-ignition-proof containers are 
the most important. Here, internal ignition is possible but 
the resulting combustion is so well controlled and contained 
that hazard is prevented. Since explosion-proof containers 
are so widely used in the mine environment, they will be 
discussed in the greatest detail in this chapter, but first, 
other, less common methods will be outlined and the topic of 
intrinsic safety will be covered. Two systems used in other 
industries, lamina and labyrinth, are variations of explosion- 
proof enclosures and will not be discussed. 

Increased safety or protection type "e," as defined by 
the International Electrotechnical Commission, is an ap- 
proach in hazard reduction where special design consider- 
ations are given to ensure an extremely low probability of 
electrical or mechanical breakdown that could produce a 
spark or temperature rise (8). As in intrinsic safety, the 
principle is to remove the ignition source. The technique is 
widely used in the Federal Republic of Germany and is 
frequently applied in mine lighting fixtures. Increased 
safety designs include such features as large spacings and 
creepage distances between live parts, protection against 
hot spots, large rotor-stator clearances for motors, superior 
amounts and quality of insulation, and special enclosures 
and fastenings to prevent unauthorized entry (8). 

Immersion systems rely on controlled environments 
to reduce or remove hazardous atmospheres from ignition 
sources. In sealing or potting methods, granular material 
such as sand is used to encapsulate the potential hazard. 
This does not actually isolate the source but quenches any 
incipient flame, thus preventing any effective contact 
between the source and the hazardous atmosphere. Purg- 
ing or pressurized systems use liquid (oil) or inert gas to 
provide protection (14). The pressurized approach is rarely 
used in mining applications (30). 

The basic problem with most of these hazard- 
reduction methods is their complexity. The overriding 
considerations for electrical hazard reduction in mining 
are robustness, interchangeability, and mobility; hence, 
there is no indication that bolted explosion-proof enclo- 
sures will lose their popularity in the near future. 

As previously defined, intrinsically safe equipment is 
incapable, under normal or abnormal conditions, of releas- 
ing sufficient energy to cause an ignition of the most 
ignitable methane-air atmosphere (39). The principle ap- 
plies to complete electrical circuits and not to individual 
components. The advantage is that safety is inherent in 
the design and is difficult to defeat. But safety can be 
compromised easily by maintenance. 

Normal conditions are taken to include the effects of 
extreme power-supply and environmental variations 
(within the equipment specifications) as well as the open- 
ing, faulting, or grounding of all conductors leading to the 



apparatus (14). Abnormal conditions involve all failures of 
internal components and wiring. 

The design of intrinsically safe circuits is a subject too 
vast in scope to cover adequately in this chapter, but some 
of the most important considerations are 

• Maintenance of suitable spacing between uninsu- 
lated conductors (only if firmly tied down); 

• Provision of additional insulation, barriers, or par- 
titions (isolation by construction); 

• The use of isolated transformer windings (for exam- 
ple, axial displacement or grounded shields between pri- 
mary and secondary windings); 

• Isolation by potting or encapsulation; 

• Limiting circuit capacitance and inductance (thus, 
the available stored energy); 

• Use of extensive circuit analysis and fault calcula- 
tions to ascertain available energy under normal and 
abnormal conditions. 

A prime concern with the last item is that with the 
failure of a single component (or subsequent failures 
resulting from the failed component) the device should 
remain intrinsically safe or fail without creating a safety 
hazard (37). The fundamental design concept usually 
applied is that two unrelated failures, each independently 
detectable but occurring simultaneously, must not com- 
promise safety. Hence, the unsafe condition could occur 
only with a third failure. Two safeguards always exist 
between the safe and unsafe conditions (14). 

References 14, 31, and 40 can be consulted if addi- 
tional information is required. Also, 30 CFR 18 contains 
extensive requirements and tests for approval of intrinsi- 
cally safe equipment (37). 



EXPLOSION-PROOF ENCLOSURES 

The explosion-proof enclosure used in U.S. mines 
complies with the applicable design requirements of 30 
CFR 18, subpart B. It is able to contain internal explosions 
of methane-air mixtures without undergoing damage or 
excessive distortion of its walls or covers, without causing 
an ignition of a surrounding methane-air mixture, and 
without discharging flame from the inside to the outside of 
the enclosure (37). Outside the United States, the same 
definition often refers to flameproof enclosures, but the 
words discharge of flame are omitted (7). Although the 
terms flameproof and explosion-proof commonly have the 
same connotation, flameproof enclosures are not presently 
allowed in U.S. underground coal mines unless they also 
meet the 30 CFR 18 requirements. 

Explosion-proof enclosures are found on all electri- 
cally powered face mining machines in U.S. coal mines 
and all motor applications in class I, division 1 locations in 
all U.S. industries. These enclosures have heavy-walled 
cast or welded construction and bolted or threaded close- 
fitting flanges. However, they are not necessarily vapor 
tight. Ignitable gas may enter a properly secured 
explosion-proof enclosure in several ways (11). Enclosures 
on machines allowed to stand for several hours in a 
gas-filled place can become completely filled with that gas 
from diffusion through openings, even though those open- 
ings are very small. Another process, breathing, results 
from the expansion of the enclosure atmosphere during 



384 



operation and contraction as it cools at rest. Air is forced 
out during expansion, and the atmosphere outside, includ- 
ing any gas present around the machine, is drawn in when 
cooling. Gas may also enter the enclosure when covers are 
removed for inspection and repair. 

A methane-air mixture in the proper proportions will 
explode if an adequate ignition source is present. The 
approximate explosibility range of methane in air is 5% to 
15^, with about 9.8*^ being the critical point for the most 
violent explosion in terms of maximum pressure, highest 
rate of pressure rise, and highest temperature. Methane- 
air mixtures above or below this range will burn but not 
explode. The lowest autoignition temperature (ignition 
without additional energy) is about 550°C (41), but a 
methane ignition requires that a sufficient gas volume is 
maintained at or above the autoignition temperature for a 
period of time. In other words, there is a critical ignition 
energy that must be injected to sustain combustion (14). 
Empirically, this minimum has been found to be 0.25 mJ 
for flammable methane-air mixtures (13). 

Ignition of the explosive gas by electrical means can 
be triggered in several ways. Arcs of sufficient energy, 
termed incendive, can result from normal or abnormal 
operation of electrical devices. Normal incendive-arcing 
modes can be attained during contact closure or even 
between fixed electrodes in a capacitive circuit, by opening 
contacts in inductive circuits, and by opening or closing 
contacts in resistive circuits (14). In the first two cases, 
energy stored in the capacitance or inductance is released 
to the arc as the contacts close or interrupt current, 
respectively (see chapter 9 on arcing and chapter 11 on 
energy storage for details). Examples of abnormal incen- 
dive arcing could involve faulted components or interma- 
chine arcing, as discussed in chapter 17. Another source of 
ignition is contact of the gas with hot surfaces, but this 
depends on the surface's heating sufficiently above the 
autoignition temperature (14). One instance of a hot 
surface would be excessive current through fine conduc- 
tors, such as in a damaged cable. 



Explosion Transmission 

Since an internal explosion is a possible occurrence, 
escaping gases generated in the explosion must not have 
sufficient energy to propagate the explosion to any haz- 
ardous atmosphere surrounding the enclosure (14). Con- 
ventionally, these hot gases are allowed to escape only 
through specially designed openings in the explosion-proof 
enclosure that provide long, narrow quenching distances. 
It is important that the gases be allowed to escape, 
otherwise internal pressures capable of rupturing the 
enclosure might develop. 

Although explosion-proof enclosures have been re- 
searched since around 1906, little concrete information 
has been gained about the phenomena that make them 
effective, but it is accepted that cooling and inhibition are 
two of the more probable mechanisms (14, 27). The hypoth- 
eses can be explained with reference to figure 16.1, which 
shows a cross-sectional sketch of a typical explosion-proof 
enclosure. An explosive methane-air mixture exists both 
inside and outside the chamber. After an explosion is 
initiated inside the enclosure, a flame front propagates 
toward the chamber walls, burning the available combus- 
tible material. This raises the temperature and pressure 
inside the chamber, and unburned gas, then heated 
burned gas, is forced through the flange gap. The jet of 
heated gas is cooled first by heat transfer within the flange 
gap. As it exits from the enclosure, it may be further cooled 
by adiabatic expansion into the surrounding atmosphere; 
during this process, combustible gas from the outside 
methane-air mixture is also entrained in the jet. Although 
the additional gas provides fuel for combustion, it further 
cools the escaping gases. If the sum of this cooling does not 
lower the jet temperature below the gas ignition temper- 
ature, the explosion will propagate to the surrounding 
atmosphere. However, if the jet entrains an excessive 
amount of combustible gas in this last phase, the heat 
supplied by the jet will be less than that lost through 
cooling, and no ignition of the outside gas will occur. 



Explosion does not propagate to 
surrounding atmosphere \ 



Unburned, then burned gases. 
No luminous flame 



Enclosure wall 
Ignition point 



Internal pressure and 
temperature raised 




Gas jet cooled by adiabatic expansion and 
entrained with surrounding atmosphere 



Gas jet first cooled by heat 
transfer in flange gap 



Enclosure cover with 
bolted plane flange 



-Gas jet 



Figure 16.1.— Cross-sectional sketch of typical explosion-proof enclosure. 



385 



The action within the flange gap is believed to go 
beyond cooling. Combustion of a gas-air mixture does not 
proceed with a single chemical reaction but rather a chain 
of reactions. If the chain carriers or active molecules from 
a preceding step are inhibited, for instance by contact with 
the flange wall, then combustion stops. This theory is also 
used to explain the protection properties of flame safety 
lamps. 

The properties of the explosion-proof enclosure that 
reduce hazards have been discussed using the flange gap as 
an illustration, but other enclosure openings such as pres- 
sure and relief vents provide the same protection phenomena 
(6). From this discussion of the theory it is obvious that the 
design criteria for joints and vents are also crucial in the 
construction of explosion-proof enclosures. 

Enclosure Joints 

Explosion-proof enclosures must have as few openings 
as possible to minimize the number of flame paths. It is 



also essential that limited access be provided for inspec- 
tion and replacement of parts. The joints formed where 
covers are placed over the required openings are either 
threaded or machine surfaced. Flanges formed by machin- 
ing may be flat, cylindrical, or a combination of both. The 
most common types of machine joints are the plane flange 
and the step flange; tongue-and-groove is uncommon. Step 
flange joints are often a combination of flat and cylindrical 
flange surfaces, as in the joint between end bells and the 
frame of explosion-proof motors. Typical cross sections of 
these joints as well as other popular types of entrances into 
enclosures are illustrated in figures 16.2 through 16.5. 
The maximum and minimum dimensional requirements 
are given in these figures and in table 16.1. Allowable 
clearances are 0.004 in for a plane flange joint and 0.006 
in for a step-flange joint in enclosures with volumes 
exceeding 124 in 3 . A simple check of the adequacy of such 
joints would be to run a 0.005- or 0.007-in feeler gauge 
(respectively) along the joint; it should not fit in the gap. 



Table 16.1. — Structural gap dimensions for explosion-proof enclosures as specified by 30 CFR 18 



Empty enclosure volume 



Type of joint 



Dimension 



Less than 45 in 3 


45 to 124 in 3 


Greater than 124 in 3 


(738 


cm 3 ) 






(2,032 


cm 3 ) 


in 


mm 


in 


mm 


in 


mm 


0.5 


12.7 


0.75 


19.0 


1.0 


25.4 


.002 


.05 


.003 


.076 


.004 


.10 


.375 


9.5 


.625 


16 


.75 


19.0 


.006 


.15 


.006 


.15 


.006 


.15 


.008 


.20 


.008 


.20 


.008 


.20 


.5 


12.7 


.75 


19.0 


1.0 


25.4 


.0015 


.04 


.002 


.05 


.003 


.076 


.5 


12.7 


.75 


19.0 


1.0 


25.4 


.003 


.076 


.004 


.10 


.005 


.13 


.5 


12.7 


.75 


19.0 


1.0 


25.4 


.010 


.25 


.0125 


.55 


.015 


.38 


.25 


7 


.25 


7 


.375 


10 



Flanged joints in 1 plane . 



Step-flange joints in 2 or more 
planes (cylinders or equivalent). 



Cylindrical joints other than shafts . 

Shafts with sleeve bearings 

Shafts with ball or roller bearings .. 



Bolts. 



Minimum width of joint 

Maximum gap (clearance) 

Minimum width of joint 1 

Maximum gap: 

Plane portion 

Portion perpendicular to plane 1 . 

Minimum length of flame path 

Maximum radial clearance 

Minimum length of flame path 

Maximum radial clearance 

Minimum length of flame path 

Maximum radial clearance 

Minimum diameter 



If perpendicular portion is more than 1/8 in but less than 1/4 in, maximum radial clearance shall not exceed 0.006 in. Neither plane nor radial portion shall 
be less than 1/8 in. 



Lock washer 



Maximum 

clearance 

0.004" 



Minimum 

stock 

Vs" 




13 capscrew 



From hole to inner edge 7/i6 n 
Minimum engagement 1/2" 



fVz-\3 capscrew 
(T ~\ ^~ Lock washer 



k-r— I 

Figure 16.2.— Typical plane-flange joint; enclosure internal 
volume larger than 124 in 3 . 



Maximum clearance 
c = 0.006" 
for i/s" < b < 1/4" , d = 0.003", 
for b>i/4", d = 0.004" 



Wi Cover i 




Shall be greater than Ve' 
a + b shall not be 
less than 3/4" 
Minimum engagement 1/2" 



Figure 16.3.— Typical step-flange joint; enclosure internal 
volume larger than 124 in 3 . 



Cover 




Hinge pin 



Total developed length to conform to 30 CFR 
18.31 (a)(b)for class Ifit 




Figure 16.4.— Threaded joint. 



Figure 16.5.— Tongue-and-groove joint. 



386 



Bolting holes must not penetrate to the interior of the 
enclosure; otherwise, the omission of a bolt would provide 
entrance to the chamber. Through-holes must be blind or 
bottomed as shown in figure 16.6. The maximum spacing 
between fasteners for joints all in one plane is 6 in. The 
maximum spacing between fasteners for joints, portions of 
which are on different planes, is based on the size and 
configuration of the enclosure, the strength of the materi- 
als, and explosion test results. The bolts or other means of 
clamping the joints should be proportioned to minimize 
stripping of threads and to give adequate strength from 
the stress developed during an internal explosion Mini- 
mum bolt diameters are also given in table 16.1. 

The minimum distance from the enclosure interior to 
the edge of the bolt hole is important as it must provide an 
adequate flame path length. The minimum Federal re- 
quirements (considering just flange width) are (37) 

Distance 
Joint size, in in mm 

V* Vs 3.2 

y 4 3 /i6 4.8 

1 or over 7 /i6 11.1 

Figure 16.3 shows how these measurements are applied. 

The Maximum Experimental Safe Gap 

To help understand the joint criteria in conjunction 
with the theory of explosion-proof enclosures, a presenta- 
tion of past research is in order. The external ignition- 
suppression properties of flange joints have been an active 
research area for several years, with extensive investiga- 
tions carried out in the United States, the United King- 
dom, and the Federal Republic of Germany. Most research 



Fuse clamp 



Fuse 



Switch-box wal 



Fuse and switch 



Machine screws holding 
switch block to casing 
wall. Hole is "blind" 



Terminal for positive 
lead conductor 




Figure 16.6.— Blind screw hole. 



work has resulted in similar conclusions, and the general 
international consensus is that design standards for 
explosion-proof enclosures are imperative. 

The maximum experimental safe gap (MESG) is a 
standard often used to determine the explosion- 
transmission properties of a flame path. The MESG is 
measured by igniting a flammable gas mixture inside a 
test system and observing if it ignites a surrounding gas 
mixture outside the enclosure (14). The MESG is the 
largest gap for a flame path length that does not permit 
ignition outside. Relatively speaking, a higher MESG 
implies a safer situation. Note that a luminous flame is 
permitted to pass through the flange, provided that it has 
insufficient energy to ignite the surrounding atmosphere. 
This does not coincide with the Federal requirements 
given earlier, since U.S. regulations do not allow passage 
of any flame (37). A gap that will not allow ignition is 7 to 
12 times larger than a gap that quenches visible flame 
(14). The following general statements can be made about 
past research findings (14, 25-26); a parenthetical com- 
ment is made where these conflict with present Federal 
regulations. 

1. The MESG increases as the enclosure volume 
decreases because thermal losses predominate in small 
enclosures. Accordingly, smaller flange widths with the 
same gap can be allowed for smaller enclosures. 

2. It is not likely that gaskets or O-rings used to make 
explosion-proof containers weathertight have any signifi- 
cant effect on enclosure safety, as long as they are external 
from the flange surface. It can be noted that O-rings are 
not permitted under 30 CFR 18 (37). 

3. There is no evidence that metallic materials are 
essential for the flange construction. (Federal regulations 
generally require metal-to-metal joints except for head- 
light and meter enclosures. Metal-to-metal flanges usu- 
ally provide a better heat-sinking effect.) 

4. Surface finish is not considered critical by most 
authorities. (Federal regulations require that flat surfaces 
be planed to within one-half the maximum clearance 
allowed, and finished to not more than 250 ^in.) 

5. The apparent safe gap between bolted flanges, the 
normal commercial construction, is considerably smaller 
than would be predicted from immovable or nonbolted 
flange tests. 

6. Evaluating the effectiveness of explosion-proof en- 
closures or studying flange-gap quenching is still highly 
empirical. 

Certain researchers have dealt with other specific 
subjects pertaining to the MESG. Torry (33), for example, 
established that no effect can be ascertained in the MESG 
over the humidity range of 0% to 50%, but the MESG 
increases as humidity increases from 50% to 100%. Con- 
sequently, the MESG can be measured with confidence 
when humidity is under 50%. By varying the methane-air 
mixtures inside and outside, in addition to in the flange 
gap, Titman and Torry (32) found the most ignitable 
outside methane-air mixture was 6.5%, with a 9.5% 
methane-air mixture inside the enclosure. James (10) 
varied the temperature in several plane-flange experi- 
ments. Using a 1.0-in (25.4-mm) joint, he found the gap 
that would just permit propagation to the outside was 
0.045 in (1.14 mm) at 80°F (25°C), and 0.035 in (0.89 mm) 
at 500°F (260°C). He stated that with a safety factor it is 
unwise to exceed a gap of 0.020 in (0.51 mm) for a 1.0-in 
flange width. Furthermore, a gap of 0.02 in or larger 



387 



cannot be applied to plane flanges without a consideration 
of enclosure material strength, since one part of the flange 
is a plate that could spring or bend during an internal 
explosion, thus increasing the gap. James also showed that 
explosion-proof enclosure safety is influenced more by 
slight changes in gaps than by changes in the flame path. 

Several researchers have discovered that doubling the 
flange width, with other conditions remaining constant, 
increases the MESG by about 1.3 (14, 25, 32). This statement 
is limited to flange widths 1.0 in and less because few data 
are available for wider flanges, which are uncommon. As the 
flange width nears 1.0 in, the relative effect of widening 
decreases. Phillips (25-26) determined quantitatively that 
the MESG should be smaller if the initial enclosure temper- 
ature is raised, and furthermore, that turbulence outside the 
enclosure increases the MESG. He substantiated empiri- 
cally the enclosure-protection theory presented earlier, and 
most of his work has been further substantiated by the 
experimental results of other researchers. 

The standards established in 30 CFR 18 (37) and by 
Underwriters' Laboratories, Inc. (36) do not permit pas- 
sage of flame and specify almost identical flange-gap and 
width values. This means that the allowable gaps found 
from MESG research are 8 to 10 times larger than the 
maximum allowable gap in the United States. Some 
foreign countries specify the gap values as a fraction of the 
MESG, and the question has been asked as to why the 
United States does not allow such larger gaps. The U.S. 
requirement can be readily justified by the need for a 
safety factor that will accommodate the problems encoun- 
tered with deterioration resulting from exposure to the 
mine environment or neglect. In fact, Short (28), in discus- 
sion with various foreign testing authorities, has found 
that none of them will accept a product with gaps as large 
as those indicated by MESG research. In practice, many 
countries permit gaps two to four times larger than those 
specified in the United States. 



Pressure vent has large effective open area for flow of gases from 
ignition but cools exiting gases and arrests flame 



Pressure vent 




Electrical enclosure walls 



Lines representing 
constant pressure 



Point of ignition 



Figure 16.7.— Pressure vent limiting pressure buildup during 
internal explosion. 



Vent body 



Swinging-door- 
vent cover 

Magnet holds ■ 
cover closed 



Bolt with / 
lock washer 




Enclosure cover 



Flame path meets MSHA 
J requirements 



-Steel retaining ring 
Vent retainer 

Flame-arresting 
vent material 
Foam stainless 
steel 



Bolt with 
lock washer 



Tapped blind hole as 
required by MSHA 



Figure 16.8.— Pressure vent assembly using metal-foam 
material. 



Pressure Vents 

Some manufacturers add special pressure-venting de- 
vices that supplement the conventional release of pressure 
through flange and shaft gaps. The concept is illustrated 
in figure 16.7 (6*). Properly sized pressure vents can reduce 
the internal pressure during explosion to 12 to 20 psi (6). 
The vents exhibit a large effective area for gases from an 
ignition but provide cooling and inhibiting to arrest the 
flame. A prototype vent developed under Bureau of Mines 
contract is shown in figure 16.8 (6). Here the escaping gas 
is filtered through flame-arresting material. 



Clamp fastened 
with bolts 

Locking screw 




Cable 



rO 



Radial clearance 
0.003" maximum 




Internally threaded 
stuffing box 

Vs" minimum (1/4" maximum) 
with cable properly packed 

F~* Hose clamp 



^O \ 



Hose conduit 



Figure 16.9.— Typical slip-fit straight stuffing box and 
packing-gland lead entrance. 



Cable Entrances 



The explosion-proof enclosure must also, obviously, 
have openings through which electrical connections can be 
made. These openings are particularly important because 
of the frequency with which trailing cables must be 
replaced on permissible mobile equipment. The most pop- 
ular type of cable entry incorporates a packing gland or 
stuffing box, which may have straight-through entry or 
angled entry. Figures 16.9 through 16.11 show cable 
entries with slip-fit stuffing boxes, where the box is a 
separate component with a cylindrical projection that fits 
into the enclosure wall (37). Specifications for enclosures 
exceeding 124 in 3 include a flame path length of 1.0 in and 
a radial clearance between the wall and box not to exceed 



Hose conduit 



Hose clamp 



Cable -*" 

Vs" minimum with cable 
properly packed 




Clamp fastened 
with bolts 



Radial clearance 
0.003" maximum 



Externally threaded 
gland nut 

Figure 16.10.— Typical slip-fit angle stuffing box and 
packing-gland lead entrance with hose clamp. 



388 



Externally threaded gland nut 

Internally threaded 
Ve" minimum -^ ".' „ ' " ' ~}~7}„ stu ^ ' ng box - ' Clamp 
clearance ^iTr II I fastened with bolts) 

Radial clearance 
0.003" maximum 

Figure 16.11.— Typical slip-fit angle stuffing box and 
packing-gland lead entrance. 




Plug shall be secured by spot welding or brazing. Weld may be on 
plug, clamp, or fastening bolt. 



0.003 in. Figure 16.12 shows how a spare cable entrance 
hole must be plugged, and figure 16.13 illustrates a 
packing gland that is an integral part of an enclosure (37). 
A stuffing box is usually constructed so that when the 
packing-gland nut is threaded into the opening and tight- 
ened, it forces the stuffing material against the cable, 
making a very tight joint. The packing material is usually 
untreated asbestos or an MSHA-accepted asbestos substi- 
tute material. When compressed, there must be at least 
V£ in of packing along the length of the cable, and the 
clearance between the packing-gland nut and the stuffing 
box must be no less than Vs in (38). 

Repacking a stuffing box in the mine can be extremely 
awkward and time consuming. A recently introduced 
method that uses a tapered polyurethane flame-resistant 
grommet in place of the asbestos (fig. 16.14), greatly 
simplifies the assembly (6). The elastomeric grommet has 
high compressibility so that one size can accommodate a 
limited range of cable sizes. 

The requirements for leads that pass between 
explosion-proof compartments separated by a common 
wall are not as rigid as those for leads passing through an 
exterior wall. One type, an insulated stud entrance, is 
shown in figure 16.15. Here the conductor does not actu- 
ally pass through the wall but is connected to a finished 
brass casting stud that is isolated from the enclosure wall 
by an insulated tube and washers. 

Windows and Lenses 

Federal regulations state that MSHA may waive 
window and lens material testing except for headlight 
lenses (37). All window and lens material must be sealed 
in place or provided with correct flange joints and must be 
protected from mechanical damage, either by guarding or 
inherently through location and structural design. If the 
exposed material area exceeds 8.0 in 2 (51.6 cm 2 ), the 
window or lens (excluding headlight lenses) must be 
protected with guarding or the equivalent. Both thermal- 
shock and impact performance tests are outlined in the 
Federal regulations. 

Enclosure Mechanical Strength and Internal 
Pressures 



/■ Plug 

/ s- Clamp (fastened with bolts) 




Radial clearance 
0.003" maximum 

Figure 16.12.— Typical plug for spare lead-entrance hole. 



Stuffing box 
(integral) part-, 
of enclosure y 



Packing 
material 



I/2" 

minimum 



i/e" minimum, 1/4" maximum clearance 
Packing -gland nut 
/ /gv* — Hose clamp 




Hose 
conduit 

Cable 



Metal 
tubing 



Locking screw 



Figure 16.13.— Typical threaded straight stuffing box and 
packing-gland lead entrance with provision for hose conduit. 



Retaining 
clip 

Cable 
entry 
body 




Tapered 
urethane 
grommet 



Figure 16.14.— Prototype trailing cable entry with 
polyurethane grommet. 



The maximum internal pressure developed during the 
ignition of an explosive air mixture in an enclosure is 
directly related to the amount of venting through flange 
gaps and any auxiliary pressure-relief devices. Tests of 
mechanical strength and internal pressures provide im- 
portant parameters used to design enclosures that will not 
transmit the explosion to a surrounding atmosphere. Mea- 
surements of explosion pressure are made to prove that the 
container strength is adequate. It is particularly impor- 
tant to ensure against flange-gap distortion. 



Insulated washers 




Insulated tube- 
Figure 16.15.— Insulated-stud lead entrance. 



389 



30 CFR 18.31(a) specifies that cast or welded enclo- 
sures must be designed to withstand an internal pressure 
of 150 psig (1.04 MPa) and that casting must be free from 
blowholes. Welded joints forming an enclosure must be 
continuous, gas-tight welds in accordance with American 
Welding Society standards. Minimum allowable thick- 
nesses for enclosure walls, flanges, and covers are also 
outlined. Again, a review of relevant research findings 
assists in understanding these regulations. 

Nagy (21) conducted extensive tests on methane-air 
and coal-dust explosions to determine pressures with vary- 
ing enclosure geometries. A range of 38 tests was per- 
formed on enclosures with internal volumes from 0.043 ft 3 
(101.5 cm 3 ) to 905 ft 3 (25.63 m 3 ), and methane-air mix- 
tures from 6.0% to 14.1%. He found that with an ambient 
internal pressure of about 14.2 psia (98 kPa), the maxi- 
mum pressure developed for all vessels was 119.6 psia 
(0.825 MPa). When the ambient internal pressure was at 
17.6 psia (121 kPa), the maximum was then 143.3 psia 
(0.988 MPa). Both maximums were obtained from exami- 
nation of a 1.0-ft 3 (2,359-cm 3 ) vessel with 9.4% methane- 
air mixtures. From 39 other experiments, using the same 
enclosure but a Pittsburgh coal-seam dust, a maximum of 
119.3 psia (0.823 MPa) occurred with a concentration of 0.6 
oz/ft 3 (601 g/m 3 ) in an 0.32-ft 3 (7.55-cm 3 ) vessel. Nagy 
concluded that 

• Vessel size does not affect the maximum pressure 
level if heat loss is neglected; 

• The rate of pressure rise decreases as the vessel size 
is increased; 

• The vessel shape does not affect the maximum 
pressure if heat loss is neglected; 

• A change in initial pressure produces a propor- 
tional change in the maximum pressure and the rate of 
pressure rise; 

• The changes in the initial temperature produce an 
inverse change in the maximum pressure and have little 
effect on maximum rate of pressure rise. 

Considering that the maximum pressure developed in an 
explosion is a function of venting, the closed-system enclo- 
sure tests carried out be Nagy could be taken as the worst 
case. The Federal regulations require enclosures to with- 
stand 164.7 psia (1.136 MPa) assuming standard atmos- 
pheric pressure, or two times the maximum pressure 
recorded during any tests without any pressure limit (37). 
The maximum permissible flange gap is rather small and 
therefore restricts venting. Consequently, the Federal 
specification does provide some margin of safety. 

Magison (14) suggests that one of the most significant 
general facts about the performance of explosion-proof 
containers is that higher maximum pressures and rates of 
pressure rise are associated with small flange gaps (for the 
same flange width with identical explosive gas mixtures). 
Referring to the work performed by the Safety in Mines 
Research Establishment (34), he states that the venting 
effect of the MESG in an 8-L sphere (488 in 3 ) reduces 
maximum pressures to a few pounds per square inch 
(ignition of 9.8% methane-air mixture). Magison indicates 
that, when considering heat loss, the actual maximum 
pressure achieved in an enclosure will depend upon the 
location of the ignition source, the container size and 
geometry, plus the arrangement of inside equipment. 
Theoretically, enclosure size has no explicit effect on the 
maximum pressure, but research has shown that size does 
have an effect for lean and rich mixtures (40). Here, the 



explanation given is that a longer reaction time occurs in 
larger vessels and allows for more heat loss. 

MSHA specifies twofold pressure tests to ascertain 
conformance with strength requirements. The first test is 
mainly to determine the explosion-proof characteristics, 
specifying that an explosive mixture within the enclosure 
will be ignited electrically and the subsequent explosive 
pressure will be recorded. The test must be repeated a 
minimum of 16 times to accommodate various ignition 
points, different atmospheric composition, and the operat- 
ing status of the motors. No part of the enclosure must 
rupture as a result of these tests, nor have permanent 
distortion exceeding 0.04 in per linear foot. If 125 psig 
(0.862 MPa) is exceeded during any test, the enclosure can 
be rejected unless it is reconstructed to reduce the pressure 
or, alternatively, is able to withstand a dynamic pressure 
twice the highest value recorded in the initial test. A 
second or static-pressure test is applied only when MSHA's 
Approval and Certification Center decides that visual 
inspection is inadequate to reveal casting or welding 
defects. The pressure applied in this test is 150 psig or 1.5 
times the maximum value recorded during the explosion 
tests, whichever is greater. 

The external surface temperature of all permissible 
mechanical or electrical components for use on electrical 
face equipment must not exceed 150°C under normal 
operating conditions in U.S. mines (37). Underwriters' 
Laboratories, Inc., has a similar specification but it ap- 
plies only to dust-proof enclosures, which will be discussed 
later. The regulation is concerned with the possibility of 
causing an ignition through surface contact between the 
enclosure and any of the four potentially hazardous mix- 
tures: methane-air, coal-dust-air, combined methane and 
coal-dust mixtures in air, and collected coal dust on the 
enclosure surface. The lowest autoignition temperature 
must obviously form the basis of the regulation. 

Of the methane-air mixtures, a 2.9%-methane mix- 
ture had the lowest autoignition temperature of about 
550°C, which can be eliminated immediately (41). Mea- 
surements to establish the minimum ignition tempera- 
tures of coal-dust clouds and dust layers have been per- 
formed by Nagy (20) on several U.S. coals. The dust cloud 
experiments on 22 different coals gave an average mini- 
mum ignition temperature of 617 °C. The lowest ignition 
temperature was 440°C for a Colorado high-volatile coal. 
Tests performed on layered dust of 16 U.S. coal samples 
showed an average minimum ignition at 222 °C, with the 
lowest temperature being 160°C for a high-volatile Illinois 
coal. Ten of the layer tests resulted in minimum ignition 
temperatures of 190°C or below. Hence, layered coal-dust 
ignition temperature is the limiting factor for the maxi- 
mum allowable surface temperature of explosion-proof 
enclosures and justifies the 150°C requirement of the 
Federal regulation. The requirement is particularly signif- 
icant for equipment in underground mines and prepara- 
tion plants where dust accumulates readily. 

Enclosure Hazards 

In recent years, some mining experts have questioned 
the inherent safety of explosion-proof enclosures. The 
concern is that when an explosion is triggered because of 
an arc or short circuit within the enclosure, gases can be 
generated that would defeat the containment properties of 
the enclosure. A number of failure modes are known to be 
possible when a high-voltage short circuit occurs within 



390 



the enclosure: these include burning through the enclo- 
sure wall, ignition transmission by hot gases or flame, 
particle ignition transmission, and enclosure deformation 
or bursting. 

The possibility exists that once an arc is established 
within an enclosure, it may jump across to the enclosure 
wall, and the intense heat and power might then cause the 
arc to burn through the wall. Killing and Tielke (12) have 
demonstrated this possibility quite convincingly. 

If organic insulation is used within the enclosure, 
combustible hydrocarbons may be present under fault 
conditions that could lead to defeat of an ignition- 
containment requirement that has been designed only for 
methane. 

Arc formation between two metal conductors (usually 
copper) within an enclosure can also cause ignition of the 
surrounding atmosphere by hot particles expelled through 
the flange gap. Heat from the arc causes metal particles to 
be separated from the conductors and subsequently forced 
through the gap by pressure buildup. These hot metal 
particles are capable of igniting a combustible gas-air 
mixture. Hence, the MESG must be small enough to 
prevent passage of such particles, further justifying the 
small gaps required by U.S. regulations. 

However, if enclosure gaps are made small enough to 
ensure that ignition of exterior gases is not caused by 
expulsion of hot gases, flame, or hot particles, the possi- 
bility then exists for high pressures to build up within the 
enclosure under fault conditions. Covers could then be 
blown off, with subsequent ignition of the surrounding 
atmosphere. Davidson and Lord (5) cited half a dozen such 
incidents in the Federal Republic of Germany, one in 
Canada, and two in the United Kingdom. The flameproof 
enclosures that burst in Canada and the United Kingdom 
had gap specifications similar to those required in the 
United States. In a typical instance, a steel cover was 
thrown several meters after shearing apart fourteen 
V^-in-diameter securing bolts. In all cases there was severe 
electrical fault damage within the enclosure, including 
charring of organic electrical insulating material. Ciok (4) 
reported that high-voltage short circuits within flameproof 
enclosures have become a matter of concern in Polish coal 
mines. Such short circuits have occurred several times, 
and the covers closing the equipment chamber have been 
blown off. 

These high pressures within the enclosures appear to 
be partly due to the evolution of gases from organic 
insulating materials incorporated in the enclosure con- 
struction. Simon (29) described experiments where organic 
insulating materials within enclosures were subjected to 
heating by an electric fault arc. He showed that in these 
tests, gases evolved that were capable of causing sufficient 
pressure to rupture the enclosure. The evolved gases 
(volatization products, consisting primarily of hydrogen, 
carbon monoxide, nitrogen, and methane) ignited sponta- 
neously upon contact with the external atmosphere. If the 
flange gaps or pressure vents would allow all the volati- 
zation products to escape, no enclosure failure would occur, 
although the nature of the escaping hot gases must be 
considered. 

Another possible contributing factor is an effect inves- 
tigated by Brown (3) and Bossert (2). It is known that 
hydrocarbon-air flames produce free ions that cause an 
electrical current to flow when the flame front bridges two 
points of opposite electrical polarity. This current may be 
sufficient to form an arc discharge; that is, an initial flame 
might produce ions that contribute to additional faults 



within the enclosure. The situation might be aggravated 
by the presence of organic insulating materials: the initial 
fault may produce hydrocarbon gases through thermal 
insulation breakdown, and the burning of these gases 
might induce additional arcing. This possibility would 
increase as voltage levels were raised, triggering a chain- 
reaction effect that could culminate in deformation or 
bursting of the enclosure. 

Materials that apparently contribute to overpressures 
are those that give off gases when subjected to heating. 
The greatest hazard is from materials that evolve combus- 
tible gases, here, insulating materials and accumulated 
water. When water is subjected to an electrical current, 
electrolysis takes place and hydrogen and oxygen are 
evolved. This constitutes a very undesirable situation; 
hence, water accumulation within enclosures should be 
avoided as much as possible, for not only does it have the 
potential of creating faults, but when a fault occurs it may 
contribute significantly to increased enclosure pressures. 

When organic insulation materials are subjected to 
electrical arcing, tracking or heating, decomposition takes 
place that yields various gases, both combustible and 
noncombustible. Over the past 30 yr, electrical equipment 
manufacturers have shown a tendency to replace tradi- 
tional insulants (cellulose, natural fabrics, asbestos, mica, 
porcelain, glass, etc.) with newly developed organic poly- 
mers. Yet at the same time, traditional insulants are still 
in common use. As a result it is possible to find almost any 
known insulation within an explosion-proof enclosure. 
The numbers are so great that it is extremely difficult to 
categorize the many types and variations of insulating and 
plastic materials that are likely to be found. However, a 
short summary of insulants that are known to evolve 
dangerous gaseous products follows (19). 

Cross-Linked Synthetic Polymers. This group contains 
the basic synthetic resins commonly utilized in the man- 
ufacture of plastic materials. Of these, the phenolic, epoxy, 
and silicone resins appear to offer the greatest hazard 
potential because of their high yield of combustible vola- 
tile products. The amino (melamine) polyester, polyure- 
thane, and isocyanate resins appear to have a lower yield 
of combustible volatile products. Materials made with 
these resins should therefore exhibit a lower, though still 
quite significant, hazard potential. 

Linear Synthetic Polymers and Elastomers. This group 
contains many of the well-known types of insulators. The 
worst potential hazards (similar in magnitude to those for 
the cross-linked group) appear to be associated with poly- 
propylene, polymethylene, polystyrene, polyethylene, neo- 
prene, and polymethylmethacrylate. Other materials 
such as nitrile butadiene (NBR) synthetic rubber, natural 
rubber, styrene butadiene (SBR) (GR S) synthetic rubber, 
Dacron polyester fiber, Teflon fluorocarbon polymer, poly- 
vinylchloride, poly ure thane, polyvinyl formal, and nylon 
show less though still very significant hazard potential. 

Other Possibly Dangerous Materials. Insulating oils 
should also be considered high hazard materials. Even 
though the gases formed are initially dissolved in the oil, 
upon saturation they may eventually evolve, causing an 
unsafe situation. A similar statement can be made about 
oil-impregnated paper. Other materials, such as cellulose 
and cotton, which do not appear dangerous at first glance, 
could also contribute to a hazardous condition because of 
the amounts of water and carbon monoxide evolved. 

Safe Materials. The electrical insulating materials 
that should be used whenever possible in explosion-proof 
enclosures are the electrical porcelains, ceramics, glasses, 



391 



asbestos, and mica. These materials are the most resistant 
to the production of gaseous products when exposed to 
heating or arcing. Note, however, that these recommenda- 
tions concerning materials consider only the possible 
contribution of evolved gases to increased enclosure pres- 
sures; toxicity is not considered. 

Although it does not appear feasible to eliminate all 
potentially hazardous materials from enclosures, the use 
of any materials that can evolve gaseous products under 
fault conditions should be avoided wherever possible. Of 
course, if a sustained arc occurs, even the conductors are 
vaporized, so all such statements are relative. Perhaps the 
best recommendation should be that the materials to be 
avoided are those that are readily volatilized, particularly 
those that evolve large amounts of combustible gases. 

PERMISSIBLE EQUIPMENT 

As defined at the beginning of the chapter, the term 
permissible equipment is applied to completely assembled 
electrical machines or components that have received 
official approval from MSHA. The term completely assem- 
bled means all equipment portions from the protection at 
the power source to all internal and external components 
of the machine, including the trailing cable. Permissibility 
requirements have been mentioned at various places in 
other chapters; details of grounding requirements were 
given in chapter 7, trailing cables and components in 
chapter 8, protective devices in chapters 9 and 10, battery 
equipment in chapter 15, and explosion-proof enclosures 
in this chapter. The aim here is to demonstrate how this 
information is tied together, by giving an overview of the 
procedures used by MSHA to investigate prospective per- 
missible equipment for safety. The overview is followed by 
a discussion of procedures recommended for checking 
equipment after it has been placed in service, and for 
maintaining explosion-proof enclosures and permissible 
equipment in proper condition. 

Permissible Equipment Schedule 

This is based on information in a 1954 Bureau of 
Mines Information Circular (9), as updated by MSHA's 
Approval and Certification Center. As already stated, the 
published regulations are contained in 30 CFR 18. Other 
pertinent regulations are found in 30 CFR 19 through 29 
for electrical equipment and 30 through 36 for mechanical 
equipment. Each of these parts is often termed a schedule. 
Schedules are revised from time to time to conform to 
equipment development and to permit as much freedom as 
possible without lowering standards. Thus, some of the 
details in the following information may become outdated, 
but the general nature of the requirements will not 
change. 

Investigations are carried out by MSHA to determine 
the permissibility of such equipment as continuous min- 
ers, shuttle cars, battery-powered vehicles, pumps, distri- 
bution boxes and so on, and for certification of components 
such as explosion-proof enclosures, connectors, and battery 
assemblies. The investigations are divided into four major 
consecutive parts: review of drawings to verify that the 
design meets the requirements, detailed inspection of the 
equipment, tests of the equipment or internal components 
in explosive gas-air mixtures and/or adequacy tests where 
appropriate, then a final inspection of the tested accesso- 
ries in the completely assembled unit or machine for 



which approval is requested. Although investigations for 
the certification of components may follow the same pro- 
cess, only the first three steps are usually necessary. 

To initiate the procedure, a written application must 
be made to MSHA, accompanied by a set of detailed 
drawings, wiring diagrams, specifications, descriptions, 
and any related material. Any intrinsically safe compo- 
nents must be stated. 

When approval is being considered, only those compo- 
nents that have a bearing on permissibility are studied; 
only one motor, controller, protective device, or unit of a 
given design is required. The investigation starts with a 
check of the drawings and specifications in order to 
determine compliance with the applicable regulations, 
and then a detailed check of all parts against the draw- 
ings, to see that they coincide. Exact measurements are 
performed on the dimensions of joints, bearings, pressure 
vents, and other possible flame-arresting paths in enclo- 
sures. For explosion-proof enclosures, for example, the 
examination determines any unnecessary through-holes, 
the adequacy of design and construction of cable and lead 
entrances, the adequacy of electrical insulation and clear- 
ances between live parts and between live parts and the 
enclosure, any weaknesses in welds or flaws in casting, 
any distortion of enclosures before tests, and the adequacy 
of the fastenings, including their size, spacing, security, 
and possible bottoming. The quality of design, material, 
and workmanship receives careful scrutiny. Only equip- 
ment adhering to the following statements will be ac- 
cepted for further investigation (37): 

1. "Electrically operated equipment intended for use 
in gassy mines shall be rugged in construction and shall 
be designed to facilitate inspection and maintenance." 

2. "Only electrical equipment that is constructed of 
suitable materials, is of good quality workmanship, based 
on sound engineering principles, and is safe for its in- 
tended use" will be tested by MSHA. 

The testing phase emphasizes explosion-proof charac- 
teristics and component properties. The general nature of 
the explosion testing has been covered earlier. No less 
than 16 internal explosions are made at various ignition 
points, using a methane-air mixture for all, with bitumi- 
nous coal dust added for some tests. The actual internal 
electrical equipment, or dummies of equal dimensions, 
must be in place for a prescribed number of tests. Motor 
rotors are tested in both stationary and rotating modes. 
An enclosure can be rejected with the occurrence of any of 
the following conditions: 

• Discharge of flame from any joint or opening, 

• Ignition of an explosive mixture surrounding the 
enclosure, 

• Development of afterburning (gas drawn into the 
enclosure by the vacuum created by the explosion, then 
ignited within the enclosure), 

• Rupture of any part of the enclosure or any panel or 
divider within the enclosure, 

• Permanent distortion of the enclosure exceeding 
0.04 in per linear foot. 

Other tests and examinations are made to determine 
the adequacy of components for the intended use. Some of 
these are performed at the discretion of MSHA investigators. 



392 



• Where the durability of a component is in doubt, 
mechanical tests will be performed to ascertain whether 
any points need to be strengthened. 

• Battery boxes are examined for ventilation, electri- 
cal clearances, insulation, drainage, and suitability for 
specific service. 

• Switches, circuit breakers, or contractors intended 
to function as switches are checked to see if they are 
capable of interrupting the maximum current permitted 
by the circuit's automatic protection device. 

• Cables, conveyor belting, and hoses are tested for 
flame resistance. 

At the end of the investigation, a final inspection is 
made of completely assembled new machines or of ma- 
chines that were previously approved but have since un- 
dergone substantial modification. The aim of the final 
inspection is to uncover any unsafe features, and the 
inspection includes such items as 

• Compliance with joint, lead-entrance, or other per- 
tinent requirements; 

• A check of wiring between components and the 
adequacy of cable clamping and mechanical protection for 
cables; survey of the positioning of cables, particularly 
those in proximity to hydraulic components; 

• Determining the adequacy of protection against 
damage to headlights, push buttons, and other vulnerable 
locations; 

• A check of the settings of overload and short-circuit 
protection; 

• Ensuring that there is a suitable means of connect- 
ing and protecting the trailing cable. 

Finally, MSHA has the option to have a staff engineer 
check the first machine produced, preferably at the factory 
where it is built. 

When approval is granted, MSHA issues a formal 
notice of approval, which is sent to the manufacturer. The 
notice is accompanied by a photograph of an approval- 
plate design. Plates reproducing the design and required 
information are mounted in a conspicuous place on the 
machine, serving to identify the machine or accessory as 
having met the applicable requirements of 30 CFR 18. 

Maintenance of Permissible Equipment 

For equipment to retain permissibility, it must be 
maintained in the same condition as that approved by 
MSHA. This is the advantage of familiarity with the 
schedule requirements: the essential details may be ex- 
tracted to establish a routine maintenance and inspection 
program that will be in full legal compliance. 

Safety can be compromised easily during mainte- 
nance procedures. One of the most common problems with 
explosion-proof enclosures is that small unapproved open- 
ings, which can constitute an immediate danger, can 
appear because of incorrect maintenance. Such openings 
are rarely caused by a manufacturing error and usually 
result from negligence or a lack of understanding on the 
part of the mine maintenance crew. One of the most 
common of such violations is the citation "open box— 
0.005," which means that a plane flange has been found to 
exceed 0.004 in (16). Other typical problems that cause 
dangerous enclosure openings are bottomed bolt and screw 
holes that have been drilled through, holes drilled through 
when plates or components have been attached, loose or 



improperly assembled cable packing glands, or undue 
wear on bearings where shafts enter an explosion-proof 
enclosure (11). 

The following precautions and procedures are recom- 
mended in the maintenance of permissible equipment (11, 
14, 16, 38), but reference should be made to the Federal 
regulations for precise compliance. Items 1 through 11 
concern explosion-proof enclosures specifically, while 
items 12 through 19 relate to permissible equipment 
covered in more detail in other chapters. Items 20 through 
23 list specific schedule requirements. 

1. Before any apparatus is examined, the power source 
should be deenergized, locked out, and tagged. This proce- 
dure is particularly important before an explosion-proof 
enclosure is opened, although it should be obvious practice 
for any electrical equipment, permissible or not. 

2. All joint clearances should be examined regularly 
with feeler gauges to determine that the maximum allow- 
able clearances given in table 16.1 are not exceeded. 

3. All pressure vents should be examined for cleanli- 
ness. 

4. Any missing bolts, lock washers, or ineffective 
fasteners must be replaced. All threaded inspection covers 
must be secured. 

5. All cover flange and thread surfaces must be 
treated with great respect. They must not be handled 
roughly, and anything, including tools, that might possi- 
bly scratch or mar a joint or thread surface must not be 
allowed to come into contact with it. If a cover thread is 
damaged, it must be replaced. If a joint is scarred by any 
means, the equipment must be removed from service. 

6. If an enclosure is opened, the cleanliness of joint 
and thread surfaces must be maintained so that foreign 
matter does not enlarge a gap. Hence, all joints and 
threads should be carefully cleaned before reassembly, and 
if necessary, a thin layer of lubricant should be applied. 
Care must be taken that the joint does not become recon- 
taminated. When reassembled, all joint clearances should 
be rechecked. 

7. Frequent examination should be made to determine 
any corrosion on flanges, threads, shafts, bearings, and 
any other flame-arresting path. Corrosion inhibitors and 
lubricants may be used, but if any corrosion is found, the 
enclosure should be taken out of service and sent for 
repairs since corrosion products cannot be removed ade- 
quately from equipment in operation. 

8. Enclosures should be examined regularly for 
burned holes. 

9. All cable packing glands should be examined to see 
that the cable fits tightly, and the clearance between the 
gland nut and housing should be checked to see that it is 
adequate. 

10. Headlights should be checked for loose or broken 
lenses, loose packing glands, missing or broken parts, and 
improper assembly. 

11. As much as practical of the accumulated coal dust 
should be removed. 

12. Portable or mobile equipment must be properly 
frame-grounded or provided with equivalent protection. 
Ground and ground-check conductors must be attached to 
separate studs that are not attached to a removable panel. 
If a separate grounding conductor is used on dc machines 
supplied by trolley power, the return (which is usually 
negative) and the grounding conductors must be attached 
to the rail or other grounded conductor with separate 
clamps. 



393 



13. All electric components must be solidly attached 
to the machine frame. Light fixtures must be grounded 
with a separate grounding conductor. 

14. No splices are allowed in the external wiring on 
permissible equipment, except for intrinsically safe circuits. 

15. All conduit hose must be flame resistant and 
MSHA accepted. The condition of mechanical cable pro- 
tection such as guards, conduit hose, and clamps should be 
examined regularly. Conduit hose should not be spliced. 
Worn or cut conduit hose may be repaired using MSHA 
approved flame-resistant cable-jacket repair material. 

16. Trailing cables should adhere to Federal regula- 
tions in type, size, length and condition (see chapter 8). A 
temporary splice must not be made within 25 ft of the 
permissible machine, except in reeled applications. Short- 
circuit protection must be provided for all ungrounded 
power conductors, either by correctly adjusted circuit 
breakers or properly sized dual-element fuses (see chapters 
9 and 10). Overload protection is recommended. 

17. All trailing cables must be provided with effective 
strain relief at the entrance to equipment. 

18. Machine cable reels and spooling devices must be 
insulated with flame-resistant material. Rollers, sheaves, 
and reel flanges must be maintained so as not to damage 
cables. Reels should maintain a positive tension on the 
trailing cable during reeling and unreeling. Reel collector 
rings should be examined for any deterioration that would 
cause a high-resistance contact. 

19. All circuit breakers and overload protection on the 
machine must be maintained in working order. A main 
circuit breaker, contactor, or disconnect switch must be 
provided on the machine and be capable of deenergizing 
all power conductors on board the machine except the 
methane-monitor power supply. Headlights and flood- 
lights must have a separate two-pole switch to deenergize 
the power conductors. 

20. All wheel-mounted equipment must be provided 
with brakes, unless the design of the driving mechanism 
prevents accidental movement when parked. 

21. If a mobile transportation unit travels faster than 
2.5 mph, headlights and red reflecting material are man- 
datory on both the front and rear of the vehicle. Such 
vehicles must have an audible warning device. 

22. Guards and safe-off devices for push buttons must 
be maintained in working condition. 

23. The approval plate must be attached to the equip- 
ment. 

Any unauthorized changes to the equipment not con- 
tained in the approval will render the equipment nonper- 
missible. Furthermore, no machine that has been changed 
from that approved may be placed in service until a field 
modification to the approval has been reviewed and ap- 
proved by MSHA. 



COAL DUST HAZARDS 

Coal dust can pose an ignition hazard of two types: a 
dust layer ignition, or dust cloud ignition. A dust cloud 
concentration above the lower explosive limit can be 
ignited by an arc, thus presenting a severe dust explosion 
hazard. The dust cloud is nonhomogeneous, and ignition is 
dependent upon the volatile content (the chemical compo- 
sition), the moisture content, and the ash content, as well 
as the size and shape of the cloud (14). When the dust is 
composed of material containing less than 8% volatile 



matter, as some anthracites and coke, there is almost no 
explosion hazard, but a fire hazard may still exist. Bitu- 
minous coals and lignite, which commonly have volatile 
contents ranging from 30% to 40%, present a serious dust 
explosion hazard. 

Bureau of Mines research (20) has determined the 
parameters required for the ignition of a bituminous coal 
dust that is minus 200 mesh dry (less than 5% moisture): 

Minimum energy mJ.. 5 

Minimum autoignition temperature ... °C. 617 

Lower explosive concentration g/L.. 0.05 

Upper limit range g/L.. 2-5 

It can be seen that the energy required for ignition is 
many times greater than that required for the ignition of 
methane-air mixtures. However, this energy level exists at 
a concentration of about 0.2 g/L and above. Concentrations 
can be measured quite readily, but the rule-of-thumb 
measure widely accepted among miners is that a dust 
cloud does not exceed an explosive concentration if a 
person can see his/her outstretched hand in front of his/her 
face (20). 

Layered dust that accumulates on equipment can also 
be a safety hazard. The parameter of most concern is the 
autoignition temperature, which depends mainly on the 
layer thickness and also the particle size related to the 
surface temperature of the equipment. Table 16.2 gives the 
minimum autoignition temperature for various thick- 
nesses of layered coal dusts (20, 24). The maximum surface 
temperature of 150° C specified in the Federal regulations 
for permissible equipment was based on this problem: A 
150° C surface temperature would allow dust thicknesses 
up to 100 mm (about V& in) without incurring autoignition, 
but at that thickness there would be almost no safety 
factor. If a 1.25 safety factor is applied to a 150° C 
equipment surface temperature, table 16.2 indicates that 
no more than 20 mm of dust should be allowed to accumu- 
late. With a 1.50 safety factor there should be no more 
than 5 mm of dust. By comparison, the Federal Republic of 
Germany requires that the maximum surface temperature 
be 75° C below the autoignition temperature for a 5-mm 
dust layer (24). The moisture level is relatively unimpor- 
tant since moisture will quickly evaporate from any sur- 
face exceeding 100° C, whether or not it is covered with 
dust. 



Table 16.2.— Minimum autoignition temperatures (in degrees 
Celsius) versus layer thickness for bituminous coals. 





Thickness, 
mm 


Low-volatile 
coal 


High-volatile 
coal 


5 




300 


240 


20 




250 


190 


50 




ND 


175 


100... 




ND 


160 










ND 


No data. 







Classification of Dust Locations 

All areas where coal dust is a hazard are defined as 
class II locations. The decision flowchart in figure 16.16 
can be used to assign the class II locations as either 
division 1 or division 2. It is based on information from the 



394 



None 



Does not exceed 
minimum depth 



Not a 

hazardous 

location 




Low 



Frequent 
or periodic 



Causes combustible 

dust cloud and 

electrical failure or 

causes electrical 

failure when 

cloud present 



Exceeds 
minimum depth 



Division 2 



Division 1 



Figure 16.16.— Decision flow chart of class II, division 1 and 
2 hazardous locations. 



NEC (37) and the literature (17, 23). Division 1 locations 
occur with 

• A low-resistivity dust, 

• A high-resistivity dust in a frequent or periodic 
combustible cloud, or 

• A plant malfunction that allows formation of a 
combustible cloud and electrical equipment to spark or 
overheat. 

Division 2 locations are defined when 

• The thickness of layered dust exceeds the minimum 
depth to propagate flame, or 

• A high-resistivity dust infrequently forms a com- 
bustible cloud. 

An additional division 2 location should be noted: when 
the layered dust interferes with heat dissipation— but 
Magison (14) suggests that this condition is implicit in 
item 1 above for any combustible dust. 

Low resistivity has been defined as values less than 
100 fi-cm; a conductive dust is one that breaks down under 
an electric field of 1,000 V/cm or less. A direct hazard 
exists with these dusts because they can provide conduc- 
tive paths between live parts and are susceptible to arc 
formation (14). Moodie (18), however, has found that coal 
dust has high resistance and is nonconductive, even when 
contaminated with mine drainage water. Hence, hazard- 
ous locations appear to exist when a cloud concentration 
exceeds 0.05 g/L or when the dust layered on electrical 
equipment exceeds a thickness of about 20 mm. These 



values have been extracted from the current literature and 
should be taken as a guide rather than a specific rule. 

Reducing Dust Hazards 

In class II hazardous locations, the types of enclosures 
in common use are dust-ignition-proof and dust-tight. 
Pressurized and intrinsically safe systems are used less 
frequently (14). The objective of these enclosures is to keep 
dust away from ignition sources and to prevent ignition of 
layered dust. Dust-ignition-proof enclosures meeting Un- 
derwriters' Laboratories, Inc., requirements conform to 
both objectives (36). The requirements are similar to those 
for explosion-proof containers but less severe. Joints, for 
example, must not be less than 3 /ie in wide, with maximum 
gaps of 0.0015 in. With wider joints, the maximum clear- 
ances are raised proportionately but must not exceed 0.008 
in. Gasketing is allowed as an alternative between mated 
surfaces, but the surface and gasket width cannot be less 
than 3 /e in. Such gaskets must not deteriorate under 
normal use and cannot be glued to the surface. The goal is 
to provide an enclosure that will prevent hot particles from 
escaping and dust from entering (14). The maximum 
allowable temperature for the surface of such equipment is 
150°C for equipment that can experience overload and 
200°C for equipment that will not usually overload. Dust- 
ignition-proof motors are suitable for both division 1 and 
division 2 locations. 

Dust-tight enclosures are intended only for division 2 
locations. The standards for their dust-tight construction 
are less restrictive than for equipment in division 1 
locations, but they undergo the same tests by the Under- 
writers' Laboratories, Inc. (35). 

Hazardous Locations in Preparation Plants 

Most of this chapter has been projected at hazard 
reduction in underground coal mines. The reason should 
be obvious: all portions of underground coal mines inby 
the last open crosscut and in return airways are considered 
hazardous locations. Hazardous locations in surface mines 
and surface facilities of mines are not so easily defined. 
The following information on coal preparation plants is 
intended to provide an example of areas that can be 
hazardous. 

Coal preparation plants commonly experience three 
conditions that are classified as hazardous: 

• High levels of coal dust suspended in the plant 
atmosphere, 

• Significant accumulation of coal dust settled on 
electrical equipment and other surfaces, 

• Dangerous accumulations of readily ignitable gases, 
such as methane released from coal being processed. 

The extent of these hazards depends on the characteristics 
of the coal being handled, the preparation plant design, 
and the steps taken to modify or control the hazardous 
condition. 

The unit operations and plant locations susceptible to 
class II hazards include (1) 

• Transfer points in the materials-handling system, 
such as conveyor-to-conveyor and bin-to-feeder locations; 

• Coal-crushing and rotary-breaking systems, includ- 
ing operations that create new particles due to coal friability; 

• Coal wetting, sizing, and sorting operations; 



395 



• Plant cleanup systems; 

• Manual picking tables; 

• Thermal dewatering systems. 

Coal dust accumulates in these locations, particularly 
inside dust covers and covered conveyors and chutes where 
effective cleanup is difficult. 

Class I locations where methane can accumulate are 
mostly operations involved in particle size reduction and 
locations that have both limited ventilation and stored or 
slow-moving coal, such as silos, storage bins, pressure 
discharge chutes, and the tunnels and chutes leading to 
these locations (1, 15). 

The purpose of the chapter has been to introduce 
hazard-reduction techniques that are used in and about 
coal mines. Portions of coal-mine operations can be ren- 
dered dangerous without diligent adherence to the proce- 
dures and regulations presented in the foregoing para- 
graphs. Anyone desirous of gaining more knowledge in 
this subject should read reference 14. 

REFERENCES 

1. Atallah, S., and P. Valence. Analysis of Coal Preparation 
Plants for Applicability of the National Electrical Code (contract 
J0166059, Arthur D. Little, Inc.). BuMines OFR 60-78, 1977; NTIS 
PB 283 411. 

2. Bossert, J. A. Electric Aging During Flammable Gas Explo- 
sions. Can. Explos. Atmos. Lab. Rep. 342, 1974. 

3. Brown, G. K., P. Morgan, and J. A. Bossert. Electric Arcing 
at High Voltage During Methane-Air Explosions. Can. Explos. At- 
mos. Lab. Rep. 293, 1973. 

4. Ciok, J. Internal Short Circuits in Flameproof High-Voltage 
Equipment. Pres. at 16th Int. Conf. on Coal Mine Safety Research, 
Washington, DC, Sept. 1975. 

5. Davidson, L., and H. Lord. Some Aspects of Fire and Explo- 
sion Hazards in Enclosure Protected Electrical Apparatus. SMRE, 
Sheffield, England, Fire Prev. and Technol., No. 10, 1973. 

6. Gunderman, R. J. Innovations for Explosion-Proof Electrical 
Enclosures. Paper in Mine Power Systems Research. 4. Transients 
and Enclosures. BuMines IC 8802, 1979. 

7. International Electrotechnical Commission (Geneva, 
Switzerland). Electrical Apparatus for Explosive Gas At- 
mospheres, Part 1: Construction and Test of Flame-Proof 
Enclosures of Electrical Apparatus. IEC Publ. 79-1, 1971. 

8. Electrical Apparatus for Explosive Gas At- 
mospheres, Part 7: Construction and Tests of Electrical Apparatus, 
Type of Protection "e." IEC Publ. 79-7, 1969. 

9. Isley, L. C, E. J. Gleim, and H. B. Brunot. Inspection and 
Testing of Mine-Type Electrical Equipment for Permissibility. 
BuMines IC 7689, 1954. 

10. James, R. S. Effect of Temperature on Flame-Arresting 
Properties of Flat Joints in Explosion-Proof Mine Equipment. 
BuMines RI 4639, 1950. 

11. Jones, D. C, M. E. Altimus, and F. N. Myers. Mechanized 
Mining Electrical Applications. PA State Univ., University Park, 
PA, Continuing Education, 3d ed., 1971. 

12. Killing, F., and M. Tielke. The Interference Short-Circuit 
With Arcing Phenomena in Electrical Equipment in Flameproof 
Enclosures With Special Reference to the So-Called Transmission 
of Ignition by Particles. Pres. at 16th Int. Conf. on Coal Mine Safe- 
ty Research, Washington, DC, Sept. 1975. 

13. Litchfield, E. L. Minimum Ignition-Energy Concept and Its 
Application to Safety Engineering. BuMines RI 5671, 1960. 

14. Magison, E. C. Electrical Instruments in Hazardous Loca- 
tions. Instrum. Soc. America, Pittsburgh, PA, 3d ed., 1978. 

15. Matta, J. E., J. C. LaScola, and F. N. Kissell. Methane 
Emissions From Gassy Coals in Storage Silos. BuMines RI 8269, 
1978. 

16. Merrihew, H. (ed.). Permissibility Checking of Electric Face 
Equipment. CONSOL, Lee Eng. Div., PA, 1976. 



17. Moodie, T. W. Explosive Dusts. Instrum. Soc. America, 
Pittsburgh, PA, Monogr. 110, 1965. 

18. Measurement Comes to Hazardous Dust Area 

Classification. Pres. at 25th Conf. Instrum. Soc. America. Instrum. 
Soc. America, Pittsburgh, PA, Paper 70-666, 1970. 

19. Morley, L. A. and F. C. Trutt. Materials Suitable for Use In- 
side Explosion-Proof Enclosures (grant G0155197, PA State 
Univ.). BuMines OFR 76-78, 1976; NTIS PB 283 972. 

20. Nagy, J., H. G. Dorsett, Jr., and A. R. Cooper. Explosibility 
of Carbonaceous Dusts. BuMines RI 6597, 1965. 

21. Nagy, J., E. C. Seiler, J. W. Conn, and H. C. Verakis. Explo- 
sion Development in Closed Vessels. BuMines RI 7507, 1971. 

22. National Electrical Manufacturers Association (New York). 
Motors and Generators. Stand. Publ. MG1. (Updated periodically). 

23. National Fire Protection Association (Quincy, MA). National 
Electrical Code. NFPA 70-1981 (ANSI Cl-1981) (Updated every 3 

yr-) 

24. Palmer, K. N. Dust Explosions and Fires. Chapman & Hall, 
London, 1971. 

25. Phillips, H. Basic Theory of Flameproof Protection. Pres. at 
IEE Electrical Safety in Hazardous Environments Conf., London, 
Mar. 16-18, 1971. 

26. Mechanisms of Flame-Proof Protection. SMRE, 

Sheffield, England, Publ. R275, 1971. 

27. The Use of a Thermal Model of Ignition To Explain 

Aspects of Flameproof Enclosures. Combust, and Flame, v, 20, 
1973. 

28. Short, W. A. Comparison of World Standards for Electrical 
Enclosures for Hazardous Area-2. IEEE Trans. Ind. and Gen. 
Appl., v. 6, July-Aug. 1970. 

29. Simon, K. F. Preliminary Investigation Into the Causes of 
the Bursting of Group I Flameproof Enclosures, in Which the Elec- 
tric Current Caused Gases To Evolve From Plastic Material. 
SMRE, Sheffield, England, Transl. 6012; 14th Int. Conf. Mine 
Safety Research Establishments, Donetsk, U.S.S.R., 1971. 

30. Stefanko, R., and L. A. Morley. Mine Electrical Systems 
Evaluation (grant G0133077, PA State Univ.). Explosion-Proofing 
of Mine Containers. BuMines OFR 76(2)-75, 1974; NTIS PB 245 
928. 

31. Thomas, V. M. Design of Intrinsically Safe Apparatus for 
Use in Coal Mines: A Review of Data and Techniques. Min. Electr. 
and Mech. Eng., v. 44, May and June 1964. 

32. Titman, J., and R. Tony. Flameproof Enclosures for Mining 
Electrical Equipment; The Protection Afforded by Flanges of One- 
Half-Inch Radial Breadth for Mixtures of Methane and Air. SMRE, 
Sheffield, England, Res. Rep. 123, 1955. 

33. Torry, R. Flameproof Enclosures for Mining Electrical 
Equipment -Influence of Atmospheric Moisture on Maximum Safe 
Gaps in Mixtures of Methane and Air. SMRE, Sheffield, England, 
Res. Rep. 202, 1962. 

34. U.K. Ministry of Fuel and Power (now U.K. Dep. Energy). 
A Review of Electrical Research and Testing With Regard to 
Flame-Proof Enclosure and Intrinsic Safety of Electrical Ap- 
paratus and Circuits. 1943. 

35. Underwriters' Laboratories, Inc. Electric Motor and 
Generators for Use in Hazardous Locations, Class II, Groups E, F, 
and G. Standards for Safety 674(A), 1973. (Updated periodically). 

36. . Industrial Control Equipment for Use in Hazardous 

Locations. Standards for Safety 698, 1974. (Updated periodically.) 

37. U.S. Code of Federal Regulations. Title 30 -Mineral 
Resources; Chapter I -Mine Safety and Health Administration, 
Department of Labor; Subchapter O-Coal Mine Health and Safe- 
ty; Part 18 -Electric Motor-Driven Equipment and Accessories; 
1981. 

38. U.S. Mine Safety and Health Administration. Coal Mine 
Safety Electrical Inspection Manual, Underground Coal Mines. 
Apr. 1979. 

39. Widginton, D. W. Some Aspects of the Design of Intrinsical- 
ly Safe Circuits. SMRE, Sheffield, England, Res. Rep. 256, 1968. 

40. Yao, C, J. deRis, S. N. Bajpai, and J. L. Buckley. Evaluation 
of Protection From Explosion Overpressure in AEC Gloveboxes 
(U.S. At. Energy Comm. contract AT(11-1)-1393). Dec. 1969. 

41. Zabetakis, M. G. Flammability Characteristics of Combusti- 
ble Gases and Vapors. BuMines B 627, 1965. 



396 



CHAPTER 17.— MAINTENANCE 1 



The ultimate goal of the power engineer is to maxi- 
mize system availability without compromising personnel 
safety. It has already been seen that protective circuitry 
with its relays, circuit breakers, and surge arrestors plays 
a vital role in achieving this goal. Yet system protection 
extends beyond protective circuitry: a good maintenance 
program can be equally important in providing a safe 
reliable mine system. Knowing the principles behind each 
type of maintenance and when and where to apply them 
can save both time and money as well as improve overall 
system safety. 

There are three types of maintenance: emergency, 
preventive, and predictive. While maintenance is a famil- 
iar topic for most engineers, this familiarity rarely ex- 
tends much beyond emergency repairs or routine mainte- 
nance. Typically, a component such as a motor fails, and in 
the heat of the production environment, maintenance 
crews make repairs as quickly as possible. Maintenance 
becomes an exercise in troubleshooting. 

Preventive maintenance (PM) is practiced by few 
mining operations, although its concept is as old as 
industrialization itself. PM is traditionally taken to be the 
periodic performance of various tasks that will extend the 
life of a component: an example is the regular lubrication 
of bearings. More generally, PM is taking measures that 
will prolong the useful life of a component, and then when 
failure is imminent, replacing it at a time when the 
minimum downtime and personnel hazard will be in- 
curred. This is the definition that will be used in this 
chapter. 

Techniques of predictive maintenance are rarely ap- 
plied in the mining industry, although the technology is 
routinely practiced in several other industries. Predictive 
maintenance is the prediction of component failure 
through measurement of key parameters and observation 
of any changes that occur over an extended period of time. 
An example is the measurement of insulation resistance 
and comparison with previously recorded values. When 
the recorded curve approaches a critical value, failure is 
impending and predictable. Advances in electronic tech- 
nology have opened the door for increasingly sophisticated 
prediction techniques, which, like the more common pre- 
dictive techniques, can be included as part of a compre- 
hensive maintenance program. 

In this chapter, various aspects of preventive and 
predictive maintenance programs are discussed, mainte- 
nance techniques are outlined, and finally, some of the 
elusive problems that can plague the mine electrical 
system are explored. But first, a few definitions are in 
order. 

The study of existing conditions in a variety of mine 
power systems indicates that a properly designed and main- 
tained power system is characterized by its safe, reliable, and 
economic operation; conversely, poorly designed or main- 
tained systems tend to be unsafe and experience low avail- 
ability (14). 2 Such conclusions are not surprising and are well 
known by power engineers working in surface transmission 
and distribution (13). What is significant is the close rela- 



1 The author wishes to thank J. L. Kohler, assistant professor of mining 
engineering, The Pennsylvania State University, who prepared the original 
material for this chapter. 

2 Italicized numbers in parentheses refer to items in the list of references 
at the end of this chapter. 



tionship between safety and availability. When the goal is to 
obtain the maximum availability of equipment, system 
safety is also maximized. 

Availability can be defined as the mean time of 
component operation, that is, the ratio of total operating 
time to the total time in which the component could have 
been operated (9). Mathematically, availability is a func- 
tion of reliability and maintainability, though it is impos- 
sible to define the precise balance of reliability to main- 
tainability that produces a specific availability. In order to 
understand and improve availability, it is necessary to 
examine the factors of reliability and maintainability 
more closely. 

Reliability can be defined as the mean time between 
failures. It is primarily affected by the component design, 
although poor reliability can be caused by abuse or im- 
proper application of equipment; for example, a portable 
cable rated for 8,000 V can be used on a 15-kV system, but 
the life and reliability of the cable will be considerably 
reduced. However, although unreliable components are 
sometimes found, and although there remain a small 
number of engineers who do not understand the principles 
of good system design, in general it can be stated that 
mine power-system reliability is essentially fixed. While 
manufacturers could try to improve the reliability of 
individual components, the results would probably not be 
commensurate with the effort and cost expended. Further- 
more, attempts to improve component performance often 
lead to increased complexity, which, paradoxically, can 
further degrade reliability. Based on this argument, it can 
be surmised that any significant improvement in the 
availability of a mine power system will have to come 
through increased maintainability. 

The U.S. Department of Defense has defined main- 
tainability as the quality of the combined features and 
characteristics of equipment design that ^permits mainte- 
nance to be accomplished under operating conditions by 
personnel of average skills (9). As implied, this factor is 
affected by the original design, but the skill and integrity 
of the maintenance personnel are also important factors in 
the maintainability of a component or of the full system. 
Another factor that is often overlooked is the ratio of PM to 
emergency maintenance. For instance, the maintainabil- 
ity of a trailing cable will be poorer if it is repaired only 
after a failure, as opposed to being repaired when signs of 
wear are discovered during a PM shift. In such a case, the 
reliability of the cable is also affected; this statement is 
substantiated by the fact that splices made on production 
shifts tend to be less reliable than those made on PM 
shifts. 3 

To increase availability, the problem is then to im- 
prove maintainability. The initial responsibility for im- 
provement would appear to lie with the manufacturer, but 
once a system is operational, manufacturer changes rarely 
impact on total availability. Consequently, the burden lies 
on the maintenance effort at the mine. Emphasis must be 
switched from repair and emergency work to PM. This of 
course is easier said than done, and experience indicates 
that PM programs at the mine are generally haphazard 
and neglected (16). 



3 Personal communication from R. H. King, Department of Mineral 
Engineering, The Pennsylvania State University, July 1980. 



397 



MINE MAINTENANCE PROGRAM 

At any mine or plant, a maintenance program must be 
justified before the company will release personnel, materi- 
als, and money. This is an unfortunate reality. The quantifi- 
cation of the costs and benefits of a PM program is always 
difficult, especially when maintenance and production de- 
partments have the misguided notion that they have oppos- 
ing goals. It is easy for a plant superintendent to visualize 
the program costs, but it is more difficult to see the long- 
range savings that will be reaped from a well-conceived PM 
program. Similarly, it is easy for a supervisor to compute the 
lost production costs when a machine is shut down for an 
hour of PM, but it is frustratingly difficult for the mainte- 
nance people to compute long-term savings due to increased 
machine availability. Unfortunately, the prevailing attitude 
among managers, from the faceboss to the superintendent, is 
that preventive maintenance is worthless unless proven 
otherwise. 

Economic Justification 

A good economic analysis is difficult to obtain if there 
are insufficient supportive data on component failure. At 
many operations, maintenance records of mine equipment 
are incomplete, at best. Consequently, it is necessary to 
make reasonable assumptions about the probable effects of 
a good PM program. Manufacturer data as well as data 
from related mine operations can be very helpful, but in 
the absence of this information there are some general 
guidelines that can be used to present a strong justifica- 
tion for a PM program. 

Three ratios are available for a first-order approxima- 
tion (25): 

cost of component repair or replacement 
cost of component PM 



cost of downtime 
cost of PM 



safety hazards 
cost of PM 



(17.2) 



(17.3) 



If the value of any one of these ratios exceeds 1.0, then PM 
can be justified. Although the application of these ratios 
should be obvious, an example of each is given. 

If a motor bearing fails, there is a good chance that 
the motor will be damaged and short out. The repair costs 
then become substantial because the motor has to be 
pulled and rebuilt. The PM costs in this case would be very 
small, consisting of monthly vibration measurements on 
the bearing. Ratio 17.1 is applicable to this case and, 
depending on the downtime involved, ratio 17.2 could also 
be used. 

The second ratio is more applicable to situations 
where the cost of the failed component is insignificant, but 
the failure causes a significant amount of downtime. For 
instance, if a contactor fails on the starter of the wound- 
rotor induction motor that drives the main conveyor belt, 
the contactor replacement costs will be minor compared 
with the cost for idling a major portion of the mine. 

Ratio 17.3 cannot be evaluated in dollars, at least not 
in good conscience. Although some industries, most nota- 
bly the automobile industry, have placed dollar values on 



human lives, this cannot be condoned. Instead, the goal 
should be to practice PM to the point of removing any 
reasonable chance of personnel injury. 

Other techniques, such as investment analysis, can be 
performed in the absence of extensive historical data. This 
method is probably best approached by considering the 
investment to be labor and overhead, and then conserva- 
tively estimating the effects of the PM program on the cost 
per unit of output. In the mine, the cost might be in dollars 
per raw ton, whereas in a preparation plant, it would be in 
dollars per clean ton. When such an estimation is carried 
out, the PM investment can be profit-adding rather than 
just profit-maintaining. However, it should also be noted 
that PM does follow the law of diminishing returns; 
carried to excess, it is not cost effective. 

Preventive Maintenance Program Implementation 

When the need for PM has been established, it is 
important that the top level of company management be 
involved in the decision to implement a suitable program. 
Since the returns from PM will initially lag behind its 
cost, such management support is crucial. Whenever com- 
pany profits fall and operating budgets must be slashed, 
the PM program is usually the first to suffer, but this is 
just the time when it is needed most. The backing of 
management can alleviate this industry-wide problem. 

The composition of the PM crew should be planned 
carefully. Experience in other industries indicates that at 
least half of the work crew should be seasoned mechanics 
who have a working knowledge of all the electromechan- 
ical system they will encounter. Circumstances may war- 
rant specialists. For example, as solid-state equipment 
becomes more prevalent in and out about the mine, an 
electronics technician would be a valuable addition to the 
PM crew. An established policy to have all PM personnel 
as salaried workers might be the most satisfactory. 

The maintenance superintendent typically assumes 
responsibility for the PM crew. However, this is not neces- 
sarily appropriate in the mining industry, where ideally 
the job should be filled by a separate individual very 
knowledgeable in electrical or mechanical engineering 
(preferably both), who has several years of mining experi- 
ence. The PM engineer should report directly to the plant 
or mine superintendent. Regardless of the crew makeup, a 
policy should be established that the PM personnel do not 
carry out normal repair maintenance. 

When implementing a PM program, an initial deci- 
sion must be made concerning the equipment to be in- 
cluded and the priorities to be followed. First priority 
should perhaps be given to new or expanding facilities. 
Maintainable equipment and installations do not just 
happen. From the planning stages through equipment 
purchasing to installation, a maintenance specialist 
should be involved in all the decisions. In fact, a recent 
study found that when this procedure was followed, the 
time required for a new preparation plant to reach capac- 
ity operation was reduced from an average of 3 yr to less 
than 18 months (23). 

When selecting the equipment for PM coverage in an 
existing operation, the components whose failure would have 
the greatest impact on production or safety should be con- 
sidered first. In an underground coal mine, these might be 
surface substations, ventilation fans and belt-drive motors, 
after which, rail-haulage motors and rectifiers, face machin- 
ery, and dewatering pump motors could be added. The 



398 



economic ratios presented earlier can be used to review the 
types of system components and assess different levels of 
coverage when designing the PM program. 

The next decision, and the one with the greatest 
impact on the success or failure of the program, concerns 
the organization of the data. The choice is basically 
between some type of card-filing system and a computer- 
ized system. Card files are rapidly becoming outdated; 
computers are becoming the normal operating method. 
Microcomputer systems now have extended capabilities 
and improved graphics and provide a convenient and 
flexible alternative that is adequate for handling the PM 
data of most companies. A wide variety of off-the-shelf 
software is available that can be readily tailored to indi- 
vidual company requirements. Such computerized card- 
filing systems have built-in analytical routines that sim- 
plify record keeping; they can print out summary reports 
of various types, print out daily calendars of work to be 
accomplished, and supplement data with graphs and 
charts. With minimal training, the PM crew can record 
data directly on to the system on a shift-by-shift, daily or 
weekly basis. 



TECHNIQUES OF PREVENTIVE MAINTENANCE 

The following sections discuss a variety of electrical 
and mechanical tests that will provide the data necessary 
for operating an effective PM program. lb these could be 
added a variety of other procedures; such simple tasks as 
measuring basic temperatures, clearances, or runout, for 
example, can provide the maintenance engineer with 
valuable information on the condition of equipment. In 
large transformers, spectroscopic analysis of the oil pro- 
vides sufficient data to detect many potential problems. 
Analysis of lubricating oils and greases can also be useful. 
The presence of a few parts per million of metal can 
indicate excessive wear or other problems. In oil-filled 
distribution transformers, for example, partial discharge 
on the windings is indicated by the presence of a few parts 
per million of copper (10). Many other techniques could be 
mentioned, but the objective of this chapter is to outline 
the measurements that would form the backbone of the 
PM program, and these are electrical and mechanical 
tests. In practice, it is important that the engineer not lose 
sight of the interrelationships between electrical and 
mechanical components in the overall system. It has been 
estimated that up to 75% of all electrical failures in motors 
can be caused by the initial failure of a nonelectrical 
component. 

Basic Electrical Measurements 

The main reason for taking measurements on the 
mine power system is to discover impending failures. 
Many potential problems can be detected with a simple 
voltmeter. Voltages on machines and at key points around 
the electrical system should be checked periodically to 
detect excessive voltage drops. Voltage drops of 20% to 30% 
under machine rated voltage are too common around 
mines, and the ramifications of such undervoltage can be 
serious, as stated in chapter 6. The voltmeter is also a 
valuable tool for troubleshooting the solid-state circuits 
that are becoming more common. 

Another useful voltmeter test is measurement of the 
voltage differences among phases at a given location, 
particularly at a transformer that supplies power to mo- 



tors. Motor damage can result if the voltage differences 
among phases are more than a few percent. In fact, one 
research group has found that the negative-sequence cur- 
rent caused by a 2% voltage difference can create enough 
heat to halve the life of a motor (25). 

The measurement of current is a simple task that can 
be performed for ac with the voltmeter and a CT or for dc 
with a resistance shunt. Current measurements taken 
periodically on a motor at no-load and full-load can provide 
valuable information on its condition. Although this type 
of measurement is better suited to stationary motors, as 
on belts or pumps, there is no reason why it cannot be used 
on mobile equipment. One specific application for current 
measurements is in the periodic checking of dual-motor 
belt drives. Here, it is normal for the load to be shared 
unequally by the two motors, for example, 52% to 48%, but 
it is not normal for the ratio to change by more than a few 
percent with time. If a significant change occurs, there 
may be an electrical or mechanical problem, such as a 
coupling misalignment. Resistance readings can also be 
invaluable. Although some of these readings must be 
made with a megohmmeter or bridge meter, many can be 
made with an inexpensive volt-ohm-milliammeter (VOM): 
resistor-bank values, open relay coils, and so on. In some 
cases, such as contact-resistance measurements, more 
accurate results can be obtained using a Kelvin bridge, 
although VOM measurements have some practical utility. 

Insulation Measurements 

Insulation is the only material used in electrical 
components that begins to deteriorate the moment it is 
manufactured. Strictly speaking, insulation should be 
called a dielectric, since it is simply one application of 
dielectric material (6). There are three primary stresses 
that can cause dielectric deterioration in the mine power 
system: mechanical, thermal, and electrical. Electrical 
stresses result from an overvoltage in excess of the rated 
withstand level of the dielectric. This may cause immedi- 
ate failure or gradual deterioration, depending on the level 
of the disturbance. 

Thermal and mechanical stresses are closely related. 
Most dielectrics do not have significant mechanical strength, 
so actual or circumferential stress may cause a failure, and 
they usually do not have much resistance to fatigue. Ther- 
mal stress, resulting from an operating temperature in 
excess of rated, causes a change in the physical properties of 
the insulator, which usually changes its electrical and me- 
chanical properties. The integrity of insulation is eroded as 
temperature increases, and several factors combine to cause 
a decrease in resistance or increase of current flow. The 
primary cause is the increased carrier mobility at elevated 
temperatures (6). Some types of insulation, especially some 
of the enamels used in motor wire, tend to become brittle, 
and as they expand and contract under the changing tem- 
perature conditions, they crack. The resulting leakage cur- 
rents will be larger than those in insulation with decreased 
resistance. 

Thermal stress can also manifest itself as a mechan- 
ical stress. This effect is more likely to occur where two or 
more dielectrics are sandwiched together, as in a portable 
cable. Since each insulator has its own particular coeffi- 
cient of thermal expansion, the expansion for each mate- 
rial will not be the same. This is particularly critical when 
an underlying material is being constrained, since me- 
chanical stresses will result. Strain may then occur, result- 
ing in voids or weakened areas. When failure eventually 






399 



takes place, it can appear to be mechanical or electrical in 
nature. 

Insulation is the material most likely to fail in a 
motor, and periodic measurements are a very valuable 
means of detecting incipient failure. The model of a 
dielectric shown in figure 17.1 indicates the parameters 
that can be measured to determine the integrity of the 
insulation (6). The two obvious ones are resistance and 
capacitance. Insulation resistance is measured directly: a 
perfect insulation would have R p = oo and R s = 0. 
Capacitance is not measured directly. Instead, 6, the phase 
angle between the voltage and the current, is measured, as 
shown in figure 17.2. The tangent of 5 (or the cotangent of 
6) is known as the dissipation factor, and the cosine of 6 is 
called the unloaded power factor (6). At power angles close 
to 90°, the two factors are nearly equal. The voltage- 
current relationships for a sound dielectric will be almost 
purely capacitive, and the unloaded power factor and the 
dissipation factor should be very low or very close to zero. 
As insulation deteriorates, large leakage currents will 
tend to increase these factors significantly (24). The dissi- 
pation factor is used frequently in technical literature, 
whereas the unloaded power factor is used more fre- 
quently by manufacturers when specifying component 
tolerances (6). Hence, the latter might be more useful in 
preventive maintenance than the dissipation factor. 

Regardless of the factor selected, readings should be 
taken at periodic intervals, and records should be kept of 
each reading so that any trends can be detected. Typically, 
a manufacturer will specify the maximum acceptable 
unloaded power factor, usually 2% or less. If this value is 
exceeded, one of two problems is usually indicated: either 



11 



R, 



ic 



R« 



Figure 17.1.— Circuit modeling a dielectric. 




I 



Figure 17.2.— Current-voltage chracteristics in a dielectric. 



the insulation is severely deteriorated or it is water 
soaked. In either case, failure may be imminent, although 
allowing the insulation to dry would correct the latter 
problem. 

A suitable instrument for making these measure- 
ments is a capacitance and dissipation-factor bridge. Since 
the temperature affects the power-factor readings, a cor- 
rection factor must be applied. These factors are depen- 
dent upon the particular dielectric type and should there- 
fore be obtained from the manufacturer. The instrument 
connections to the apparatus being tested are identical to 
those of the megohmmeter tests, which are discussed in 
the next section. 

The shunt resistance of the leaky capacitor, R p , shown 
in the insulation model of figure 17.1, can be measured 
and used to predict insulation failure as well as to observe 
some measure of the insulation integrity. In practice, this 
measurement is usually performed with a megohmmeter. 
The current that flows because of the potential impressed 
on the dielectric is composed of two parts: a leakage 
component across the insulation surface and a current 
actually through the insulation. When the detection of 
deterioration is of interest, the superficial leakage current 
must be minimal, otherwise the insulation resistance 
reading will be artificially low. Consequently, it is neces- 
sary to examine those factors that can cause incorrect 
measurements. They are 

1. Surface conditions: abrasion, foreign material, 
moisture; 

2. Moisture; 

3. Temperature; 

4. Method of test: test-instrument potential, duration 
of test; 

5. Residual charge. 

The surface condition of the insulation normally has 
the largest effect on the superficial current component; for 
instance, an abraded surface will collect airborne dust. If 
the dust is from a highly conductive material such as from 
carbon brushes, the possibility of establishing a conduc- 
tive path on the insulation surface is rather high. Even 
dust with a much lower conductivity (such as coal dust) 
can allow enough current to flow so a significant error is 
introduced. A nonabraded surface will have a tendency to 
collect some dust, but here the problem is not as severe. 

A conductive path can also be created if the dielectric 
surface is moist, and foreign material on the surface will 
aggravate the problem considerably. The surface moisture 
problem can usually be minimized by taking the measure- 
ments only when the temperature of the component is 
above the dewpoint. In the case of machines, surface 
moisture is of no significance at the normal operating 
temperature, but some researchers have expressed concern 
over the effect of high relative humidity during testing (4). 
While it is true that experiments conducted in a controlled 
environment do reveal a decreased insulation resistance 
with increased relative humidity, the changes are not 
troublesome if the data are correctly analyzed. More will 
be said about this later. 

The second factor is moisture other than surface 
moisture. Some insulating materials are hygroscopic; they 
absorb moisture. This effect is entirely different from that 
of surface moisture and presents a more serious problem 
because it compromises the dielectric integrity. Although 
a dielectric may exhibit a greater propensity for absorbing 
water as it ages, this characteristic is not a reliable 



400 



method of determining aging, at least not outside the 
laboratory. 

The hygroscopic moisture content can be reduced by 
bringing the component to normal operating temperature 
for a few hours before making any measurements. If a 
measurement is low and moisture absorption is suspected, 
caution should be observed when energizing the compo- 
nent, to prevent failure. In some cases, such as with large 
motors, it may be worthwhile passing a current equal to 
load current through the windings at a low voltage. The 
heating that occurs will dry out the windings. 

The third factor, insulation temperature, has a signif- 
icant effect on the measured resistance. If the insulation 
temperature is known, a normalized resistance can be 
calculated: 



R c = R t K t 



(17.4) 



where R c = insulation resistance corrected to 40° C, Q, 

R t = measured insulation resistance, fi, 
and K t = temperature coefficient. 

The temperature coefficient factor is obtained from a 
graph such as that shown in figure 17.3, which is valid for 
rotating machines (11). 

The measurement of insulation temperature is some- 
times difficult, but neglecting this factor can be risky. A 
suitable compromise can be reached in many cases; for 
example, taking the measurements consistently before or 
after the component is at operating temperature may 
stabilize the readings. The ambient temperature may vary 
over a small enough range to ignore its effect on the 
preoperating temperature of the component. Of course, 
steps such as these must be evaluated for each situation, 
and what might work well for a large motor could be 
meaningless in the case of a circuit breaker. 

The fourth factor, method of test, is important but 
presents no real application problems. The test potential 
of a megohmmeter is usually 500 to 1,000 V. Measure- 
ments on a given component should always be taken at the 
same test potential. The basic test format for a given 
component is fixed and will be suggested shortly. Beyond 
this, the test measures and procedures should always be 
duplicated precisely. Since the purpose of the test is to 
analyze a trend over time, small errors introduced by 
nuances of the test procedure will not pose a problem, 
provided that the methodology is consistent. 

Residual charge is the final factor. The leaky capaci- 
tor of the dielectric model must be discharged before a 
resistance measurement is taken. This is achieved by 
shorting to ground the outer skin of the dielectric and the 
enclosed conductor. The short should remain in place for a 
few minutes. When multiple readings are taken on the 
same insulation system, the dielectric should be shorted 
for at least 3 min, assuming the standard 3-min test (11). 
In specialized megohmmeter tests, where parameters such 
as the time and rate-of-rise of resistance are analyzed, this 
factor is critical. 

Megohmmeter Tests 

As was described in chapter 5, a megohmmeter is a 
portable instrument with an integral voltage source and a 
meter calibrated in megohms. The voltage source is either 
a handcrank generator or a battery pack. The type of 
instrument most suitable for routine insulation resistance 
measurements has only two probes, much like a voltmeter. 



Tests are conducted by placing the probes across the 
insulation system and measuring the resistance. Compo- 
nent manufacturers generally provide instructions for 
megohmmeter testing of their components. Three differ- 
ent tests are performed: spot reading, time resistance, and 
multi voltage. 

Spot Testing 

Readings for the spot test are taken consistently at 
60s intervals, since the resistance in a good dry insulation 
will always increase with time, as shown in figure 17.4. 



-z. 

UJ 

o 

u_ 
u_ 

Ld 

o 
o 

UJ 

en 
cr 

UJ 

a. 

UJ 



o 

-z. 
< 

H 
CO 

CO 

UJ 

tr 




0.05 h 

i i i i i i i i 

20 40 60 80 

WINDING TEMPERATURE, °C 

Figure 17.3.— Graph relating approximate insulation 
resistance variation with temperature for rotating machines. 




4 6 8 10 

TIME, min 

Figure 17.4.— Insulation resistance versus application time 
of test voltage. 



401 



Spot readings are taken periodically on system compo- 
nents and recorded so that any persistent downward 
trends that indicate potential failure can be detected. 

Conceptually, the most straightforward spot test is on 
cables, where one power conductor is checked at a time, 
with all other cable components shorted together and to 
ground. The connections for checking the line A conductor 
of a cable are shown in figure 17.5. After the connections 
are made, the test voltage is applied for 1 min and the 
final resistance value is recorded. 

Megohmmeter testing of motors is similar, but this 
test checks only the insulation system to ground and not 
the turn-to-turn insulation system. Figure 17.6 shows the 
test connections for line A testing of an ac motor, and 
figure 17.7 shows the dc motor connections. The minimum 
value for the spot resistance test should always be 1.0 MQ, 
plus 1.0 Mfi for each 1.0 kV of nominal voltage rating for 
the equipment. Consequently a low-voltage mine motor 
should have a minimum resistance of 2.0 MO. Typical spot 
resistance test records for motors are shown in figures 17.8 
through 17.10. 



Disconnect other 
end of cable 



Any grounding 
conductor or shield 




"X 



Megohmmeter 



Figure 17.5.— Megohmmeter test connections for checking 
cable insulation in line A. 



Brush 




Megohmmeter 
Figure 17.7.— Megohmmeter test connections for dc motor. 



100 



c| 50 



UJ 

o 

11 

CO 
UJ 

or 



10 



~i ■ r 



20 40 60 80 100 120 140 

TIME, months 

Figure 17.8.— Spot resistance curve for normal motor. 




Megohmmeter 



3 



UJ 

u 

-z. 
< 
I- 
<£> 

CO 

UJ 

or 



t i i i I t~~ i r~ 




1 I i i i i i_ 



Figure 17.6.— Megohmmeter test connections for ac motor. 



INCREMENTAL TIME — 

Figure 17.9.— Spot resistance curve showing effects of dust 
and moisture. 



402 




158 162 
TIME, months 

Figure 17.10.— Spot resistance curve for defective motor. 



When spot-testing a transformer, the core iron must 
be grounded and the resistance connections or solidly 
grounded connections must be removed. All windings 
except the one under test must then be shorted together 
and grounded. The test connections are shown in figure 
17.11. Similar testing procedures exist for starters, control 
boxes, relays, and circuit breakers, to name a few. 

Time-Resistance Tests 

With the same connections, time-resistance tests can 
be performed. Here successive readings are taken at 
specific time intervals to form time-resistance curves as 
shown in figure 17.12. Of particular interest is the point 
where the curve begins to level out, since a good insulation 
will have a continual increase in resistance with time (26). 
Figure 17.13 shows curves for an actual deteriorating 
motor. 

The polarization factor can be calculated by plotting 
the resistance values at 1 and 10 min: 



r _ R io 

Ri 



(17.5) 



The computed factor should have a minimum value of 2.0, 
as shown in figure 17.14. The polarization factor is inde- 
pendent of temperature and equipment size, which makes 
it very convenient for some applications. A polarization 
factor curve for a deteriorating motor is shown in figure 
17.15. 

Multivoltage Tests 

Multivoltage tests are particularly valuable for as- 
sessing the efficiency of high-voltage components. Insula- 
tion resistance measurements are repeated at different 




Megohmmeter 



Grounding 
resistor 



Figure 17.11.— Megohmmeter test connections for 
transformer. 







TIME 
Figure 17.12.— Time-resistance curve. 



UJ 

o 



oo 

CO 
UJ 




Figure 17.13. — Three time-resistance curves for 
deteriorating motor. 



403 



100 




120 



Figure 17.14.— Time-resistance curves showing polarization 
for hypothetical motor. 




77 80 83 86 89 92 95 98 101 104 107 110 
TIME, months 

Figure 17.15.— Polarization factor curve for deteriorating 
motor. 



voltage levels, such as 100, 250, 500, and 1,000 V, and the 
resistance found at the higher voltage should always be 
equal to or greater than the resistance at the lower 
voltage. A large reduction in insulation resistance with 
measured voltage is indicative of insulation weakness, as 
shown in the curve for a deteriorating motor in figure 
17.16 (26). 

Harmonic Analysis 

As insulation deteriorates, a small initial current that 
is rich in harmonics can be produced from electrical 
discharge across the voids (26). The mechanism producing 
the harmonics is probably partial discharge, which will be 
discussed later in this chapter. The presence of harmonic 
current provides a method of detecting insulation failure, 




6 
TIME, min 



Figure 17.16.— Multiple voltage curves for deteriorating 
motor. 



since researchers have shown that in new insulation the 
frequency component of a current is primarily that of the 
source, whereas in old insulation the current is composed 
of many frequencies (26). 

Harmonic analysis is carried out by a spectrum ana- 
lyzer, which performs a Fourier analysis on a signal and 
displays the amplitude of each frequency component. An 
instrumentation arrangement to carry out the analysis is 
shown in figure 17.17, where current is flowing from a 
60-Hz source. The CRT display of the spectrum analyzer 
immediately indicates insulation of questionable integ- 
rity. The method has proved to be effective for cables, 
motors, and transformer-insulating oils. 

Power Factor Versus Voltage 

In this test the unloaded power factor is plotted 
against various voltage levels. For perfect insulators the 
resulting curve is a straight line, but in deteriorating 
dielectrics the curve rises with increased voltage to form a 
parabola, as shown in figure 17.18 (26). This is because 
voids in insulation contain entrapped air, which tends to 
ionize as voltage approaches the breakdown point, thus 
producing the extra loss that causes divergence of the 
plotted curve. A description of a simple power factor meter 
for in-mine measurements can be found in reference 26. 

Infrared Testing 

Although it is not an electrical test directly, the 
observation of infrared emissions with a portable instru- 
ment can detect abnormal hot spots in operating electrical 



404 




To 

unloaded 

system 



High-voltage 
transformer 

Figure 17.17.— Circuit for harmonic tests. 



01 

o 

h- 
o 

UJ 

o 

Q_ 



1 


Coil with 
/ many voids 




/ .. Coil with 
/ few voids 


>^--- 


1 Tip-up 

i 
i 
i 
i 
— '■ 1» 



VOLTAGE, kV 

Figure 17.18.— Power-factor versus voltage curves showing 
tip-up. 



equipment. At any temperature above absolute zero, all 
bodies radiate energy. The radiation in the infrared region 
is closely proportional to the body temperature. Excessive 
temperature can be caused by broken conductor strands or 
excessive insulation leakage in cables, defective coils in 
motors, or excessive leakage flux in transformers (26). 
When using routine infrared observations as part of the 
maintenance plan, any sudden temperature increase can 
indicate trouble. 

Real-Time Computer Analysis 

The two greatest problems concerning PM in the 
mining industry are probably the need to depend on 
personnel to make the necessary measurements and the 
inadequacy of existing techniques for predicting many of 
the failures that occur. The search for ways to alleviate 
these problems has been an ongoing process (18, 20). One 
recently developed method, real-time computer analysis, 
eliminates the human link in the data collection and 
analysis process. The method is also sensitive to deterio- 
ration processes that cannot be detected by more conven- 
tional tests. The technique is based upon the classification 
of a deterioration matrix composed of the values of elec- 
trical parameters, such as power-frequency harmonics and 
symmetrical components, that are collected continuously 
in a real-time environment. Economic analyses indicate 



that the cost of such a system could be amortized within 2 
yr. However, practical implementation of the method is for 
the future. In the meantime, much can be gained through 
using the conventional tests described here. 



MECHANICAL MEASUREMENTS 

The failure of simple mechanical devices on a ma- 
chine, such as a bearing, can lead to a catastrophic 
electrical failure. Electrical problems can cause excessive 
vibration (as with open-end rings on rotor-bar circuits), 
which in turn can cause electrical deterioration as com- 
mutator damage, insulation cracking, and squirrel-cage 
rotor damage. In fact, a variety of electrical problems can 
mimic mechanical deterioration. The only way to pinpoint 
these things is to include mechanical measurements in 
the maintenance program. Indeed, in the case of rotating 
machines, it is the most applicable method. 

Vibration 

Machine vibration is an important symptom of elec- 
trical or mechanical problems. Electrical problems that 
can cause vibration include the presence of a subharmonic 
voltage component (15, 30, or 45 Hz) or shorted turns in 
one phase of a motor. Vibration can also be caused by such 
mechanical problems as unbalance, misalignment, or 
faulty parts. 

Unbalance can occur in a motor shaft or coupling as 
the result of improper installation, mishandling or defect. 
It may be static, occurring in a single plane, or dynamic, 
occurring in more than one plane. Whatever the cause, the 
vibration should be eliminated to prevent commutator 
damage, winding-lead fatigue or cracking of insulation. 
Some devices, such as fans, require balancing both before 
use and at periodic intervals. Coupler misalignment can 
also lead to vibration problems, especially when the motor 
coupling to a pump or air compressor is not in perfect 
alignment. Faulty parts can cause vibration. For example, 
when a bearing begins to fail, its vibration level will 
increase. If vibration levels are recorded routinely, for 
example, every 60 days, any significant increase will be 
detected, incipient failure can be predicted, and problems 
avoided. 

The measurement of vibration is not too difficult. The 
transducer used is typically a linear accelerometer, which 
for convenience is often attached to a magnet or probe, as 
shown in figure 17.19. The output signal is connected to 
an instrument that gives a readout of acceleration, veloc- 
ity (by integrating the acceleration signal), and displace- 
ment (the second integral of the acceleration signal). 
These instruments are also available with an oscilloscope 
display for examining the measurements in real time and 
in time or frequency domains. 

Although analysis of vibration data is beyond the 
scope of this chapter, a few comments are in order. First, 
taking readings involves engineering judgment. In the 
rather simple case of the motor and pump, readings would 
be taken along each axis at points A through D, as shown 
in figure 17.20. Careful analysis is required to avoid 
various pitfalls; here for example, an unwary reader might 
confuse cavitation with vibration. Table 17.1 gives a brief 
summary of some of the more common causes of vibration, 
and figure 17.21 provides some insight into the severity of 
different vibration levels (27). 



405 



Table 17.1.— Common causes of vibration 



Cause 



Amplitude 



Frequency 



Phase 



Remarks 



Unbalance Proportional to 

unbalance. Largest in 
radial direction. 
Misalignment of couplings Large in axial direction, 
or bearings and bent 50% or more of radial 

shaft. vibration. 



Bad bearings — antifriction Unsteady— use velocity Very high, several x 
type. measurement if possible, r/min . 

Eccentric journal Usually not large 1 x r/min 



Bad gears or gear noise... Low-use velocity measure Very high, gear teeth x 

if possible. r/min. 

Mechanical looseness No data 2 x r/min 

Bad drive belt Erratic or pulsing 1,2, 3, and 4 x r/min of 

belts. 

Electrical Disappears when power 1 x r/min or 1 or 2 x 

is turned off. synchronous frequency. 



1 x r/min Single reference mark . 



Most common cause of vibration. 



Aerodynamic hydraulic No data 1 x r/min or number of 

forces. blades on fan or impeller 

x r/min. 

Reciprocating forces do 1 , 2 and higher orders x 

r/min. 



1 x r/min usual; 2 and 3 Single, double, or triple.... Best found by appearance of large axial 
x r/min sometimes. vibration. Use dial indicators or other 

method for positive diagnosis. If sleeve 
bearing machine and no coupling 
alignment, balance the rotor. 

Erratic Bearing responsible, most likely the one 

nearest point of largest high-frequency 
vibration. 

Single mark If on gears, largest vibration in line with 

gear centers. If on motor or generator, 
vibration disappears when power is 
turned off. If on pump or blower, attempt 
to balance. 

Erratic No data. 

2 reference marks. Usually accompanied by unbalance and/or 

Slightly erratic. misalignment. 

1 or 2 depending on Strobe light best tool to freeze faulty belt. 

frequency. Usually 

unsteady. 
Single or rotating double If vibration amplitude drops off instantly 

mark. when power is turned off, cause is 

electrical. 

No data Rare as a cause of trouble except in cases 

of resonance. 



..do. 



Inherent in reciprocating machines; can 
only be reduced by design changes or 
isolation. 



Magnetic probe BNC conne ction 



Magnet 




Magnetic probe 



Probe extension- 



Keeper (remove before 
placing magnet on machine) 

BNC connection 




Cable 



Figure 17.19.— Mounting techniques for two vibration 
transducers. 



Vertical Motor Shaft and Motor load 
, T j coupling / " 

Axia' 



Horizontal 



Figure 17.20.— Four typical vibration measurement points. 




E .E 



Z -O 



t_> 01 

< 3 

_1 u> 

Q_ O 

to <y 



a. 



0.05 - 



0.01 r 



0.005 - 



0.001 



VIBRATION FREQUENCY, Hz 

1,000 5,000 50,000 

500 | 2,000 I 10,000 | 100,000 




1,200 | 3,600 
1,800 

Figure 17.21.— Typical vibration severity chart. 



Q 



< 

Q: 
m 

> 



406 



Acoustic Emission 

All rotating machines produce wide-band acoustic 
emissions that can provide considerable insight into the 
operating efficiency of a component. In fact, acoustic- 
emission analysis can predict most incipient mechanical 
failure, ranging from shaft defects and bearing problems 
to bad welds, many days in advance of the failure. It is a 
relatively new technology that has not yet been widely 
adopted in the mining industry. Portable instruments are 
available for measuring and displaying acoustic signals, 
using either contact or noncontact probes at specific loca- 
tions on the machine. Figure 17.22 illustrates the tech- 
nique (12). 

The distributions of greatest interest are the fre- 
quency and the amplitude, and both play a role in emis- 
sion analysis. A crack propagating in a bearing, for 
example, generates a narrow acoustic-emission pulse with 
a flat frequency distribution that is almost identical to 
friction noise in the bearing assembly. In this case, the 
frequency emission alerts the engineer to a potential 
problem. Subsequent analysis of the amplitude distribu- 
tion will identify the specific problem, since the emission 
pattern for a propagating crack and that caused by friction 
are readily distinguished. 

Acoustic emissions divide conveniently into low- 
frequency emissions, ranging from about 5 to 1,000 Hz, 
and high-frequency emissions that start around 65 kHz 
and rise into the megahertz range, but 1,000 kHz is 
considered the practical cutoff point. Although low- 
frequency signals can be used to diagnose many machine 
malfunctions, they have the disadvantage of containing 
many frequencies caused by normal phenomena, and these 
complicate interpretation of the readings (12). High- 
frequency emissions do not have this problem with back- 
ground noise, and furthermore, they are more defect 
oriented. The signals in the high-frequency range are 
rapidly attenuated, which makes it easier to locate the 
source of the emission. 

CONTINUOUS-MONITORING SYSTEMS 



Low-frequency vibration spectral data 





Envelope-detected high-frequency spectral data 

1 — ■ — i — ■- 

Defective 
Good 
bearing 




300 4000 100 

FREQUENCY, Hz 



200 300 400 



Figure 17.22.— Comparison of acoustic-emission tech- 
niques for detecting failing roller bearing. 



Mine 
environment 




Sensors 




Communication 
link 








































Figure 17.23.— Conceptual diagram of generalized mine 
monitoring and control system. 



Recent research in predictive techniques has focused 
on the development of reliable continuous-monitoring sys- 
tems that will automatically identify disturbances in 
patterns of normal operation and thus give warning of 
impending problems. The systems employ contact and 
noncontact sensors located at strategic locations on equip- 
ment and linked via a communication network to a central 
processing facility. Research designers favor distributed 
systems with outstations or local data processors that 
collect and display information and serve as a filter, 
passing only important data through to the central sta- 
tion. The outstations may have limited decision capability 
and may trigger actuators that are part of a system control 
network. These features are diagrammed in figure 17.23. 
Remote sensing and control systems have the potential to 
reduce downtime, thereby increasing productivity, and it 
is anticipated that the use of such systems will become 
routine in mines. 

Esoteric spectral analysis and remote sensing and 
control can seem far from the day-to-day routines of the 
electrical maintenance engineer who is concerned with 
basic repairs and troubleshooting defective equipment. 
Few tasks can be more daunting to the young, inexperi- 
enced engineer, armed only with theoretical knowledge, 
than being faced with a downed system while surrounded 



by an impatient production crew. Troubleshooting is a 
skill— some would claim an art— that is acquired through 
experience. Many of the techniques outlined earlier in this 
chapter are employed to diagnose equipment problems, 
and these together with troubleshooting tips and layouts 
provided by equipment manufacturers are usually suffi- 
cient to identify the cause of the trouble. There are, 
however, some instances where the problems are more 
puzzling and defy routine analysis; among these are 
corona and partial discharge. These topics will be dis- 
cussed in the remainder of this chapter. 

CORONA 

Corona is the name given to very small transient 
discharges that occur as part of the process of localized 
gaseous ionization associated with dielectric materials. 
Corona is most prevalent in high-voltage systems, and 
when the process reaches critical levels in regions of high 
electrical stress, the byproducts of ionization can cause 
degradation of insulation and lead to system failures. 
Awareness of the problem became widespread in the 
mining industry in the early 1970's when 12.47-kV distri- 
bution systems were adopted in several underground 



407 



mines. The increase in system integrity expected to result 
from the high voltage did not materialize; instead, the 
change was accompanied by anomalous failures in cou- 
plers, cables, and stress cones. Ensuing research added 
greatly to the understanding of partial-discharge phenom- 
ena and the practical implications for mine power systems. 
Corona is now recognized as a concomitant feature of high 
voltages. In this section the conditions for the inception of 
corona, its subsequent behavior, and the effects on compo- 
nents in the power system are outlined. 

Figure 17.24 shows a general graph of conduction 
effects in a gas such as air. These effects can be explained 
in terms of the ionic theory of conduction (28, 30). Up to 
point A in the graph, Ohm's law is valid; at point A, 
saturation occurs because of the space-charge density. 
After the potential is increased to point B, the gas begins 
to ionize and the field changes from a subdischarge field to 
an ionizing field. With any further increase in potential 
there will be complete breakdown; in other words, a 
discharge field is attained. 

The proportionality between current and voltage up to 
point A occurs because the electric field is so weak that it 
ionizes very few gas molecules. Hence, the number of ions 
and free electrons is very small compared with the number 
of gas molecules. Multiplication effects are minimal, so in 
general the current depends only on the mean speed of the 
ions and their relative numbers. The formation of ions and 
free electrons is approximately equal to the number of 
recombinations, and the current will depend on the rate of 
this progress. Saturation between points A and B occurs 
when the electron density reaches such a level as to 
decrease the field intensity; the flux of ions and electrons, 
thus current, remains constant. A further increase in 
potential causes the ions and free electrons to acquire 
sufficient velocity that, when they collide with neutral 
molecules, enough energy is imparted to split the mole- 
cules into ions and free electrons. This process increases 
exponentially; thus, the current also increases exponen- 
tially. This ionization by collision is the most important 
mechanism of conductivity in gases. 

The most complete breakdown in the process of ion- 
ization by collision gives rise to a spark. Essentially, a 
spark consists of a quantum of electrical energy traveling 
through the gas, with the associated current limited only 
by the source. It is very unstable because the passage of 
current lowers the voltage. Below complete breakdown, 
there are three other types of discharge that are known 
collectively and rather loosely as corona. These are the 
discharges of interest here. In descending order of magni- 
tude they are: 

• Brush discharge: These discharges are often consid- 
ered an anomaly of corona and consist of a very small 
number of sparks ending in air. 

• Corona discharge: The onset of this discharge is 
signaled by a glow or halo ("corona" is the Latin word 
meaning "crown"). 

• Partial discharge: This discharge cannot be de- 
tected visually; chemical changes occur in the gas. 




VOLTAGE 
Figure 17.24.— Conduction in gas. 



Ld 
CD 
< 



O 
> 



Ionization 
by collision 



C0r0na RriKh 

discharge J£SL 

Partial , discharge 

discharge ., 



1 



Spark 



Arc 



CURRENT 
Figure 17.25.— Discharge sequence in an ionizing field. 



for completeness, but this discharge is not a product of 
ionization by collision; it is an unstable condition that 
occurs when the electrodes are hot enough to supply 
electrons for the current, thus vaporize the anode, and 
create positive particles that heat the cathode by impact 
(see chapter 9). 

Any system where corona may be initiated is known 
as a corona source. The specific area of discharge, the 
corona site, is the source of transient currents that pulsate 
for a few nanoseconds with a current too small to be 
measured directly. When the corona pulses occur at regu- 
lar intervals for several minutes, the phenomenon is 
known as continuous corona; where the periods succeed 
each other at increasing intervals, it is known as intermit- 
tent corona. The frequency may vary from 1 pulse per 
minute for dc, to 100,000 pulses per second for 60-Hz ac 
(8). The shape of the corona pulse varies widely, being 
affected by the corona current at the site. 

Discharges are a response to electrical stress. Mathe- 
matically, the conditions for the occurrence and mainte- 
nance of the discharge are given by Townsend's continuity 
theorem (28): 



In this chapter the term corona will be used to describe 
the phenomena of partial, corona, and brush discharges. 
Figure 17.25 is a plot of the ionizing field for gases 
showing the different types of discharges. In any given 
situation, the sequence of discharges will always progress 
in the order shown, but the discharges can begin at any 
point in the sequence (28). The arc in the figure is included 



[ ' a[exp ( X (j8 - a)dx]dx = 1, (17.6) 



where a = ionizing coefficient of positive ions, 
j8 = ionizing coefficient of negative ions, 
and I = path of ionization (e.g., wire). 



408 



A very important part of Townsend's theory shows that 
discharges occur only in nonuniform electric fields; these 
are also the points of greatest electrical stress concentra- 
tion. Common geometric configurations conducive to high 
stresses and thus to the formation of corona are shown in 
figure 17.26. 

Corona Behavior 

The inception and behavior of corona from an ac 
potential can be illustrated by considering a single bare 
conductor that is parallel to a ground plane. In this case, 
the dielectric or insulator is air. A high-voltage test set is 
connected to the conductor, and as the voltage of the 
conductor is increased, the gradient between the conduc- 
tor and the ground plane rises. When the gradient reaches 
a critical value, the air molecules at the conductor inter- 
face are ionized. The voltage at this point is referred to as 
the corona inception level. 

The ionization rate is a function of the air tempera- 
ture, pressure, and the potential gradient. Higher gradi- 
ents result in higher ion and electron velocities, hence, 
greater energies. When these particles strike other mole- 
cules, the energy transfer is sufficient to cause ionization, 
and there will be an exponential increase in ionized gas. 
The ionization process will continue outward until the 
ionized particles no longer have sufficient energy to split 
any other molecules. During this process, the conductor 
will have the characteristic violet halo of corona discharge, 
extending outward for two or three times the conductor 
diameter. 

Closer examination of the conductor will reveal a 
group of bright beads, evenly spaced and superimposed on 
the otherwise uniform halo. The beads are negative co- 
rona, which occur on the negative side of the sign wave; 
the uniform glow is positive corona, which occurs on the 
positive half. The inception level for negative corona is 
higher than that for positive corona and negative corona is 
believed to be more destructive (22). Sensitive detection 
equipment has been used to determine the macroscopic 
ionization frequency, which varies from 1 to 10,000 Hz (8). 

If the line voltage is reduced to the corona inception 
level once more, the corona will be significantly reduced 
but it will not disappear. In the case of the bare conductor, 
a surface discharge, the corona extinction level is slightly 
below the inception level. For a discharge in a void, the 
extinction level is 15% to 20% less (24). 

In the literature, particularly the early literature, 
researchers tended to be indiscriminate in their use of 
corona terminology. Two terms are commonplace: thresh- 
old corona level, meaning the point at which it occurs, and 
visual corona level, meaning the point at which it can be 
seen. Since corona is now detected and quantified using 
sophisticated instrumentation that includes ultraviolet 
light imperceptible to the human eye, the interpretation of 
visual corona level is subject to confusion, and great care 
must be exercised when comparing data given in the 
literature. 

The characteristics of dc corona are basically the same 
as those of ac corona, for their respective polarities. 
However, dc corona usually occurs intermittently and at a 
slightly higher potential. Since the dc conductor will not 
change polarity, unlike the ac conductor, a surface charge 
is deposited by the initial discharge, which must leak 
away before another discharge can occur. This decreases 
the frequency of discharges. The ac corona is often consid- 







A Corner 




B Point 



Conductor 



C Interface 




Insulation 



D Abrupt change in electric E Irregular symmetry 

field 

Figure 17.26.— High-stress geometries. 



ered to be the mean of positive dc and negative dc coronas 
(15). 

Harmonics due to corona can range into several thou- 
sand, but the third harmonic is the chief development (21 ). 
There is a dual relationship between harmonics and 
corona: harmonics are caused by corona, and there are also 
corona losses associated with the harmonics. In transmis- 
sion systems, the latter are the most important. 

In 60-Hz systems, corona occurs during part of each 
half cycle, and the corona discharge is pulsated at twice 
the line frequency. This causes a cyclic change in line 
admittance, which results in a modified or distorted wave- 
form (6). The sinusoid has considerable harmonic content, 
particularly the third harmonic; the fifth, seventh, and 
ninth harmonics are often present. Fourier analysis of the 
symmetrical components of the current shows that the 
third harmonic must be composed only of zero-sequence 
currents or voltage; therefore, if the lines are connected in 
a grounded-wye configuration, the triple-frequency 
(third-harmonic) currents caused by corona will flow 
through the lines and into the ground loop. Zero-sequence 
currents do not flow in delta-connected systems; instead, a 
triple-frequency voltage pulsates between lines. The ap- 
pearance of other harmonics as currents or voltages is 
easily determined by knowing the sequence of the har- 
monic in question (18). 

Ionization during corona causes radiation of radio 
frequency noise. Below the visual corona level, radio 
frequency noise is negligible, on the order of 10 /xV. The 
radio interference voltage increases rapidly up to 100 or 
200 /xV with the occurrence of visual corona (17). 

When air is ionized during a surface discharge, ozone 
and nitrogen oxides are the principal byproducts. If mois- 
ture is present, the nitrogen oxides combine with the 
water to form nitric acid, which causes gradual deteriora- 
tion of the insulation. Ozone is very unstable and changes 
quite rapidly to harmless diatomic oxygen, but a thin 
layer of ozone persists in the vicinity of the insulation. 
This can cause the insulation to harden, to become brittle; 
and if the site is under flexure stress, it will develop large 
cracks. 

The effects of corona in small or microscopic voids 
within insulation can cause more serious damage because 
the site is not visible to external inspection. Virtually all 
commercially available insulated conductors contain air 
within small random dielectric voids. Such voids usually 






409 



occur as bubbles with diameters ranging from 0.1 to 0.01 
mm. Despite excellent quality control, present economics 
and technology prohibit the manufacture of insulated 
conductors that do not contain some voids. An air film may 
also exist between the conductor-semiconductor interface, 
or between layers of insulation, and occluded gases have 
also been observed in the interstices of the dielectric, as 
shown in figure 17.27 {24). Voids can also result from 
careless handling or poor splicing practices in the mine. 

Little is known about the air spaces within a given 
insulation. The size, location, distribution, pressure, and 
gas content of the voids are all unknown, making corona 
evaluation difficult. The boundary between the dielectric 
void and the insulation can be represented as two materi- 
als having dielectric constants E x and E 2 with a potential 
across them (fig. 17.260 (24). This series combination of 
the two materials results in an effective dielectric con- 
stant, E k , which is less than either E x or E 2 . From 
electromagnetic theory, it is known that this effect is due 
to the increased electric-filled intensity at the interface, 
caused by the discrete change of the dielectric property (7). 
The breakdown potential of the contained air is lower 
than that of surrounding air; thus in a void, corona can 
occur at a voltage that is below the rated corona extinction 
level. 

Degradation of insulation by corona can be explained 
by two separate mechanisms that result in the chemical 
decomposition of the dielectric. Electron bombardment is 
believed to be the primary mechanism; chemical degrada- 
tion is a secondary mechanism. As explained earlier, 
ionization by collision depends upon the continuation of 
energy transferal during collision. Depending on the po- 
tential gradient, the ionization extends outward for some 
finite distance. However, in a dielectric void, the contained 
gas will be ionized, and then the ions and free electrons 
strike the molecular structure of the dielectric. This en- 
ergy transfer is usually sufficient to break the weak bonds 
and produce volatile products with a lower molecular 
weight (5). As the void begins to deteriorate, the electrical 
stresses increase, causing a snowball effect, until the 
surface of the insulation is pierced. Depending on the 
location of the punchthrough and the specific cable appli- 
cation, a line-to-line or a line-to-ground fault may occur. 
The corona site may not result in an immediate failure but 
nevertheless remain active, continuing to deteriorate the 
dielectric material. 

During surface ionization, the chemical products of 
the process are solely responsible for dielectric damage, 
but in voids chemical degradation can be considered a 
secondary mechanism. Volatile byproducts of ionization, 
which include some very caustic acids, enhance the dete- 
rioration process during ion bombardment. 

Corona Detection 

Numerous tests have been devised to determine the 
presence of corona, but the tests have been designed 
primarily for checking corona in commercially produced 
cables and, hence, are of limited use in the mine. Never- 
theless, the principles underlying these devices could be 
applied with slight modification for mine use. Since the 
current of the corona pulse is too small to be measured, its 
effect on a traveling wave is monitored. The change in the 
traveling wave can be related to the apparent charge of the 
corona pulse, which is proportional to the damage caused 
by the corona. An apparent charge greater than 4 pC is 
usually considered harmful. 



A basic system for detecting corona discharge (fig. 
17.28) consists of a partial discharge detector and display, 
power-separation filter, power supply and voltmeter, and 
high-voltage transformer. The power supply is used to 
deliver a voltage high enough to initiate corona. The 60-Hz 
signal and its harmonics are filtered out using the power- 
separation filter, leaving only the high frequency due to 
corona discharge. The detector contains the electronics 
necessary to integrate the pulses and determine the peak 
pulse values. Various circuits are used to perform these 
operations. Practically all detectors work on this principle, 
though variations are sometimes incorporated to obtain 
results under special or adverse conditions. The technol- 
ogy in this area is still changing rapidly. 

A fundamentally different technique that uses ultra- 
sonics has possible application to mine power systems. 
This method is based on the fact that corona breakdowns 
cause both audible and ultrasonic pressure waves at the 




A Missing semiconductor tape 

B Void created by conductor damage 

C Metal particle with associated air pocket 

D Bubble 

E Blister in splice jacket 

Figure 17.27.— Typical dielectric voids in cables. 



Partial-discharge detector 
and display 

-f- 



Power supply and voltmeter 



High-voltage transformer 




Figure 17.28.— Block diagram for corona-detection system. 



410 



corona source. If the medium containing the corona source 
is in free-moving air, these pressure waves can be detected 
by an ultrasonic transducer. Successful tests using a 
barium titanate transducer have been reported (2). The 
advantages of this method are the portability of the 
equipment and the relative simplicity with which the 
presence of discharges is determined. Disadvantages in- 
clude the need to add complex equipment to obtain quan- 
titative data. When traveling through a solid, the pressure 
wave will follow the path of least resistance, thereby 
increasing the time of travel and making it difficult to 
determine the exact location of the corona site. An addi- 
tional problem is that noise in transformer cores, known 
as magnetostriction noise, renders the ultrasonic detector 
useless around transformers. 

Partial-Discharge Problems in Mining 

In mine power systems, some corona destruction goes 
undetected until failure occurs. Subsequent analysis of 
failed cables, couplers, or stress cones frequently identifies 
the culprit as partial discharge, the lowest level of dis- 
charges associated with corona. In other cases, partial 
discharge is the suspected cause but a lack of corroborative 
data precludes a direct correlation. Instead, such failures 
can be attributed to anything from "bad cable" to "tran- 
sients," which may not be entirely accurate. 

In high-voltage systems that require a 15-kV insula- 
tion class, problems with partial discharge are common in 
cables and couplers. Most transients do not have sufficient 
energy to cause failure in good cable, but when a cable has 
been weakened by partial-discharge degradation, it be- 
comes susceptible to failure or at least to a higher rate of 
deterioration because of the increased stress. Ultimate 
failure can come from a single large energy transient or 
from a number of smaller transients over a period of time; 
hence, it is important to minimize transients in the mine 
power system. 

Partial discharge can be initiated by improper cable 
handling, particularly bending the cable sharply or pass- 
ing cable directly over metal parts or through insulation. 
Any abrupt change in the electric field along the cable 
length causes sufficient stress to initiate partial dis- 
charge. The subsequent chemical degradation can eventu- 
ally reach the conductor surface and terminate with a 
fault. Cable applications should therefore be investigated 
carefully before they are installed in power centers, distri- 
bution transformers, or other units. A specific installation 
can be partial-discharge resistant in one case, but suscep- 
tible in another. 

The severe discontinuity in the electric field that can 
occur at the termination of a high-voltage cable is a prime 
site for partial discharge, and various stress-relief systems 
are employed to prevent its inception. Some of these 
methods are shown in figure 17.29 and were discussed in 
detail in chapter 8. Similar problems are found in the 
confined spaces of high-voltage couplers, where stress 
relief is provided by preformed filler moldings or stress- 
control tape. With both methods, it is extremely important 
to eliminate air voids, since partial-discharge site can be 
created by any manufacturing or mounting defect. 

Couplers can be subjected to other damage caused by 
partial-discharge byproducts as the result of careless han- 
dling in the mine. A coupler with a folded insulator tube, 
for example, can create a restricted area with lower 
surface resistivity at the fold. Ozone and nitric acid can 



Copper braid 




Semiconductive 
tape 



Semiconductive 
tape 




Molded cone used with 
semiconductive tape and 
copper braid 



Cone molded with 
stress-control semi- 
conductive tape 



Stress-control 
tape 




Heat-shrink 

stress-control 

sleeve 




Stress-control tape 

Figure 17.29.— High-voltage cable terminations 



Stress-control 
heat- shrink 






form at this site and cause rapid deterioration of insula- 
tion inside the coupler housing. 

Recent research into discharge phenomena within 
high-voltage couplers identified typical sites for the forma- 
tion of partial discharge and demonstrated the importance 
of adopting impeccable standards of workmanship when 
repairing or reassembling couplers (19). It was shown that 
the smallest void, cavity, or discontinuity can become a 
partial-discharge site; examples are 

• Small nicks or cuts in cable insulation. 

• A void at the termination of extruded semiconduct- 
ing tape. 

• A service loop in the grounding strap that is too 
long and is located too close to the uninsulated power 
conductors when assembled within the shell. 

• An uninsulated void between the pin and the orig- 
inal cable insulation (fig. 17.30). 

• Nuts on conductor pins that, when turning, have 
caused conductor strands to untwist, forming voids. 

• Voids in insulation caused when pinned conductors 
are inserted into insulator tubes and the middle section of 
the coupler shell. (A well terminated coupler end might 
have discharge levels below 3 pC but may increase to 30 
pC or more when inserted into the insulator tubes.) 

• Voids at the bottom of insulator tubes (fig. 17.31). 

• Voids formed when potting compounds cure and 
shrink away from the coupler shell. 

The use of a detector such as that diagrammed in figure 
17.28 was found to be essential when attempting to isolate 
partial-discharge problems in failed couplers. 

The extinction level of partial discharge is an impor- 
tant parameter in all parts of the mine power system. The 
critical value for initiating or extinguishing partial dis- 
charge is the potential to the ground plane. Since the 
neutral in high-voltage mine distribution systems estab- 
lishes the ground plane, the main concern is line-to- 
neutral voltage. Obviously, the extinction level must be 
safely above the nominal line-to-neutral system voltage; 



411 




Figure 17.30.— Major insulation void sometimes found in 
high-voltage coupler terminations. 




Shell 



Figure 17.31.— Possible stress site in high-voltage coupler 
insulators. 



otherwise, partial discharge initiated perhaps by a tran- 
sient overvoltage would continue. A safety factor of 25% 
above the maximum line-to-neutral voltage is consistent 
with power-engineering practice (3). This converts to an 
11-kV extinction level for the 15-kV couplers used on mine 
distribution systems. 

In general, the recommendations for reducing partial 
discharge in the mine system are 

• Minimize high-voltage transients. 

• When installing cables inside power centers, belt 
transformers, and so on, take care not to bend or stretch the 
cables unless they have been designed for that purpose. 

• Pay special attention to eliminating insulation 
voids and possible partial-discharge sites in high-voltage 
systems. 

INTERMACHINE ARCING 

Intermachine arcing, another maintenance problem, 
refers to electric arcing between the frames of under- 
ground electrical face equipment that is of sufficient 
magnitude to ignite explosive methane-air mixtures. Wolf 
(29) has described the sources and corrections for ac 
intermachine arcing on mobile face equipment, and this 
section summarizes his work. 



When the 1969 Coal Mine Health and Safety Act 
specified that all low-voltage and medium-voltage trailing 
cables for ac mobile equipment must contain an insulated 
conductor for the ground-monitoring circuit, mine opera- 
tors could not comply immediately because available ca- 
bles did not meet the specified requirements. The demand 
prompted conversion from three-conductor type G round 
and flat cables to a newly developed type G-GC. In this 
cable, one of the grounding conductors is insulated to serve 
as a pilot conductor, and the size of the remaining ground- 
ing conductors is increased. The result is a cable with 
asymmetrical cross section. In 1971, mining companies 
began replacing their mining equipment trailing cables 
with this new type, but soon after, sparks were observed 
arcing between equipment frames. Following exhaustive 
tests by the Mine Enforcement and Safety Administration 
(MESA, now MSHA), it was concluded that this sparking 
was caused by an induced voltage in the grounding con- 
ductors of continuous miner trailing cables that originated 
from the cable asymmetry. Subsequent investigations con- 
firmed that the energy released by the arc was incendive. 

A review of the conditions associated with this phe- 
nomenon indicates that the problem increases with 
trailing-cable length, the power demand of the machine, 
and the asymmetrical geometry of the cable. The presence 
of the insulated ground-check conductor in the cable 
causes a physical asymmetry that aggravates the problem. 
Such is true whether the cable is shielded or nonshielded. 

There are four basic methods for solving the induced- 
voltage problem and the subsequent arcs, which will be 
discussed in turn: 

1. Transposition of the phase conductors, 

2. Use of symmetrical-cable types, 

3. Use of diodes to suppress arc currents, and 

4. Use of saturable reactors to suppress arc currents. 

Transposition means to change the power-conductor 
position with respect to the individual neutral conductors 
at specific locations along the cable length. Induced volt- 
ages in the grounding conductors can be cancelled by 
dividing the cable precisely into thirds, and cutting and 
resplicing it as shown in figure 17.32. This method is 
based on the fact that if a grounding conductor is located 
the same distance from the three line conductors, the 
voltages induced by the line currents sum vectorially to 
zero. Equal transposition of the power conductors along 
the cable length has the same effect. 

Although this method is both simple and effective, its 
disadvantages inhibit its use in mining. It shortens cable 
life and physical strength, and in actual practice equal 
transposition lengths are difficult to maintain, and per- 
haps most important, correct transposition can be easily 
lost during subsequent splicing. 

The use of symmetrical cables is an alternative solu- 
tion. In theory, symmetrical cables exhibit no induced- 
voltage or arcing-current phenomena. The common sym- 
metrical cables are the type G round and type G + GC 
round (see chapter 8). However, a symmetrical cable that 
has been spooled or one that has been physically abused 
exhibits considerable distortion and is not truly symmet- 
rical. Further, any unbalanced power-conductor load or 
conductor defect will alter the cable symmetry, and the 
resulting induced voltage will be impressed on the ma- 
chine frame. Despite these problems, the use of symmet- 
rical cables appears to be the best long-term solution to 
the problem of intermachine arcing. It should be noted 



412 



that induced grounding-conductor voltages can be negated 
only if the continuity of each individual neutral conductor 
is assured through monitoring. This is obviously impossi- 
ble with SH-D cable types. 

Another approach to solving the problem is to insert a 
nonlinear impedance in the grounding circuit. The device 
must have a high effective impedance at low induced 
voltages, but a low effective impedance to the higher 
voltages and currents that are available during a ground 
fault. Any such device must have a continuous-current 
rating of at least 25 A and have a short-circuit capability 
equivalent to the cable grounding conductors. 

Diodes connected in a bridge arrangement, as shown 
in figure 17.33, apply this concept. The center diodes 
conduct at all times while the other diodes conduct every 
alternate half-cycle. The devices that are usually used 
conduct about 1.0 A with a voltage drop of 0.6 V. Since an 
incendive arc occurs with about 1.0 A, the danger of 
explosion can be eliminated with a sufficient number of 
diodes. For instance, with an induced voltage of 6.0 V, 10 
diodes in series are needed, which means that a total of 12 
diodes are required in the bridge arrangement. Because 
simultaneous line-to-neutral faults on different lines at 
different system points are possible, if the diodes fail, it is 
essential that they fail short, not open. 

A saturable reactor is a nonlinear element, which can 
limit arcing current without affecting fault current. The 
saturable reactor has several advantages over diodes, 
mainly, that there is no junction of semiconducting mate- 
rial subject to damage. The device is simply an iron-core 
inductor that has a high impedance to the point where the 
core is saturated. The inductance of the reactor is much 
less after saturation than before. This means that the 
device will limit induced voltages and arcing currents to a 
safe value, while operating as a linear device. For fault 
currents and voltages, the saturable reactor operates in 
both linear and nonlinear regions; in other words, the 
currents are not sinusoidal, but the effective impedance 
will be low because the greater part of operation is in the 
nonlinear region. Thus, the grounding system is still 
effective. 

The selection of a suitable saturable-reactor charac- 
teristic is somewhat similar to that for diodes. The main 
input is the amount of induced voltage present. Once this 
is known, a reactor can be chosen that will operate in a 
linear region for the induced voltage, limiting arc current 
to 1.0 A or less. The reactor should saturate for a slightly 
higher voltage and thus be effective in limiting machine 
frame potentials during faults. An acceptable saturable- 
reactor characteristic is shown in figure 17.34. 

Saturable reactors can be damaged by high separate- 
line, simultaneous-fault currents, and the winding conduc- 
tor can be fused by the large current if its capacity is not 
great enough. As with diodes, the device must fail short in 
these cases. Another problem is that saturable reactors 
can store energy, which lowers the atual incendive cur- 
rent. Such energy storage can be restricted by selecting a 
low quality factor (X/R ratio). 

Saturable reactors and diodes are equally effective if 
they are installed between the machine frame and the 
grounding conductor or in the power center between the 
grounding conductor and frame ground. But when they 
are installed in power centers, the grounding pin in 
metal-enclosed cable couplers must be isolated from 
ground; otherwise, the device will be shorted out. (This is 
the same precaution as that recommended in chapter 9 for 
wireless ground-check monitors.) 



'/3- point 



2 /3- point 





I 
Red 


I i 
! Red 


! Red 


Ground i \ 


71 !\ 


■J' 


White ! 


\ ! White ! 


K ! White 


Ground ! V 


: iV 




Black \/ y 


\ i Black |/ > 


\ Black 


Ground 















Splice 



Splice 



Figure 17.32.— Power-conductor transposition on three- 
conductor type G cable. 




■ Power center 



Figure 17.33.— Application of diode-suppression bridges in 
power center. 



'rms 



10 
5 



•rms 



5 10 



Figure 17.34.— Typical saturable-reactor characteristic. 



GROUND DIRECT CURRENT OFFSETS 

The dc offsets in grounding can be a serious problem 
wherever mixed ac-dc systems are used. Intermachine 
arcing can result from stray dc as well as from ac induced 
voltage on grounding conductors. The presence of dc 
ground currents on the ac grounding system can also cause 



413 



protective-relaying malfunctions with ground-fault relays 
and ground-check monitors. This section discusses the 
methods available to eliminate these offset currents, while 
information on the sources of the currents can be found in 
chapter 7. 

A discussion of corrective methods cannot be re- 
stricted to dc offset currents alone since they are only one 
anomaly caused by dc superimposed on the ac ground in a 
bipower ac-dc system. All major problems, including 
ground-bed deterioration by electrolysis, must be consid- 
ered. The analysis can be conveniently divided into three 

parts: 

» 

1. Feasible power-system modifications, 

2. Grounding-system construction, and 

3. Water-distribution systems. 

Power-System Modifications 

The dc finds its way onto the ac ground by direct 
contact between equipment frames and the mine floor and 
also through two-wire dc shuttle car supplies where the 
negative lead is tied directly to the frame of the shuttle car 
and rectifier. In the first case, stray current from the 
mine's trolley system flows through the mine floor, roof or 
ribs, then moves through an ac equipment frame and into 
the ac safety-ground circuit. Stray current tends to flow in 
the mine floor because here it finds a lower resistance path 
back to the rectifier through the track. The resistance is 
usually higher in the track because it is difficult to bond 
track systems satisfactorily underground. Some bonding 
procedures are hard to follow and lead to improper bonds, 
while other methods are easier to use but result in a bond 
that is physically unsound and easily damaged. Even if a 
good bond is achieved, it can be destroyed by heavy rail 
usage or even the type of abuse caused when moving 
continuous miners into the section over the tracks. 

A parallel feeder, bonded at each rectifier and also 
bonded to the track at 100-ft intervals, should greatly 
attenuate the amplitude of stray dc flowing in the mine 
floor. Stray dc from trolley sources should present little 
problem when this arrangement is used in conjunction 
with careful ac equipment and rectifier placement, for 
example, positioning equipment a minimum of 25 ft from 
the track or, if this is impossible, placing ac equipment on 
insulating mats. 

The use of isolated or floating dc feed for shuttle cars 
would help to alleviate the problem with shuttle car loads. 
A two-wire negative-ground system is cheaper than a 
three-wire arrangement, but this represents a serious 
electrical compromise, as does diode grounding. A two- 
conductor type G shuttle car trailing cable is advocated to 
prevent offset problems. Although it is not in practice, 
ground-check monitoring of the dc cables is worthwhile. 

Grounding-System Construction 

The dc offset problems are minimized with correctly 
constructed grounding systems. Construction of a ground 
mat or grid is governed by three needs: to achieve a low 
value of earth resistance, to control potential gradients, 
and to prevent corrosion. Low earth resistance is achieved 
by placing a sufficient length or surface area of metal in 
intimate contact with the soil. If the soil has a high 
conductivity, less metal is required than if soil resistivity 
is high. For safe grounding, a very low resistance is 



required between the soil and the buried metallic grid. 
Potential gradients are controlled by the depth of burial 
beneath the surface and the placement of the grounding 
conductors in relation to one another. The manner in 
which a grounding grid responds to the flow of current 
through it depends on the magnitude and duration of the 
loading. Generally, all metal structures should be tied 
together to eliminate potential-gradient hazards. How- 
ever, corrosion protection necessitates the isolation of 
underground metallic structures from the corrosive effects 
of the soil. Detailed information on these subjects is 
provided in chapter 7. 

Water-Distribution Systems 

There is no doubt that underground water-pipe sys- 
tems can serve as excellent sinks for any stray currents 
moving through the mine floor. Water pipes should be 
isolated from current sources as much as possible. Where 
stray currents are present, nonconductive pipe sections 
could be inserted in the waterline at intervals, which 
would serve to limit the amount of current carried by the 
pipe. However, a problem remains because of the water 
inside the pipe. Although distilled water is an insulator, 
the presence of dissolved ions renders it conductive, and 
most mine water is rich in impurities. Hence, the best 
protection is to maintain isolation (as much as practical) of 
the water-distribution system from possible stray ground- 
current sources. 

The suggestions given here on dc offset ground cur- 
rents must be general in nature because no two mine 
systems have identical problems. Hence, guidelines are 
more effective than specific rulings in this instance. When 
this material is combined with that in chapter 7, a 
thorough list of correction methods for reducing the effects 
of ground dc offset currents can be assembled. 

SUMMARY 

The foregoing sections in this chapter have covered 
many aspects of maintenance, including justification, 
measurements, and planning, as well as some specific 
problems related to the subject. Many other areas fall 
under this general title and its supervision in the mining 
industry. These aspects are major portions of chapters 6 
through 16, where they are presented in detail. Using this 
knowledge as background, the main objective of chapter 
17 has been to integrate additional information into a 
concept for a feasible PM program. 

The importance of adequate PM cannot be overempha- 
sized. Through its employment, equipment availability 
can be maximized while personnel hazards are minimized. 
On the other hand, poorly done PM is worse than worth- 
less, and managers who have experienced this are very 
opposed to any approach to PM in their mines. 

Chapter 17 brings to a conclusion this publication. The 
objective has been to assemble a comprehensive engineering 
reference on mine power systems. Because not all aspects of 
mine electrical systems could be included and because this 
area of mining is rapidly changing and growing, the aim has 
been to collect as much significant information as possible 
that will provide the basic tools needed to continue a knowl- 
edgeable involvement in mine electrical applications. Such 
an involvement is important today, but in the future it will 
have even greater significance as the use of sophisticated 
electrical systems expands. 



414 



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21. Peek, F. W. Voltage and Current Harmonics Caused by Cor- 
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22. Peek, F. W., Jr. Dielectric Phenomena in High Voltage 
Engineering. McGraw-Hill, 1920. 

23. Rushton, J. N. Maintenance Considerations in the Design of 
a New Plant. Min. Eng., v. 30, Mar. 1978. 

24. Shrader, J. E. Corona in Air Spaces in a Dielectric. Trans. 
Am. Inst. Electr. Eng., v. 41, Sept. 1922. 

25. Smeaton, R. W. (ed.). Motor Application and Maintenance 
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26. Stanek, E. K, M. M. Hassan, Y. C. Chou, and H. Shamash. 
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27. Vibranalysis Engineering Corp. (Houston, TX). Short 
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28. Whitehead, S. Dielectric Phenomena. Van Nostrand, 1927. 

29. Wolf, R. A., E. K. Stanek, and P. Kantabutra. Intermachine 
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8694, 1975. 

30. Woodruff, L. F. Principles of Electric Power Transmission. 
Wiley, 1938. 



415 



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416 



APPENDIX.— ABBREVIATIONS AND SYMBOLS 



UNIT OF MEASURE ABBREVIATIONS 1 



A 


ampere 


Ah 


ampere-hour 


A/m 


ampere per meter 


A/m 2 


ampere per square meter 


A//is 


ampere per microsecond 


As 


ampere-second 


A/s 


ampere per second 


A/V 


ampere per volt 


AAVb 


ampere per weber 


C 


coulomb 


°C 


degree Celsius 


cm 


centimeter 


cm 3 


cubic centimeter 


C/m 2 


coulomb per square meter 


cmil 


circular mil 


C/s 


couloumb per second 


C/V 


coulomb per volt 


°C/W 


degree Celsius per watt 


deg 


degree 


F 


farad 


°F 


degree Fahrenheit 


F/m 


farad per meter 


ft 


foot 


ft 3 


cubic foot 


ft?/h 


cubic foot per hour 


ft- lb 


foot pound 


ft«lb/V 


foot pound per volt 


ft 3 /min 


cubic foot per minute 


ft//ts 


foot per microsecond 


g 


gram 


g/L 


gram per liter 


g/m 3 


gram per cubic meter 


H 


henry 


h 


hour 


H" 1 


recriprocal henry 


H/m 


henry per meter 


hp 


horsepower 


Hz 


hertz 


in 


inch 


in 2 


square inch 


in 3 


cubic inch 


in/ft 


inch per foot 


in/s 


inch per second 


J 


joule 


J/K 


joule per kelvin 


J/s 


joule per second 


K 


kelvin 


kA 


kiloampere 


kg 


kilogram 


kg/m 3 


kilogram per cubic meter 


kn 


kilohm 


kn/v 


kilohm per volt 


kPa 


kilopascal 


kV 


kilovolt 


kVA 


kilovoltampere 


kVA/hp 


kilovoltampere per horsepower 


kvar 


kilovar 



1 Standard IEEE format (rather than Bureau of Mines format) is used for 
unit of measure abbreviations in this publication. 



kV/ft 


kilovolt per foot 


kV/ M s 


kilovolt per microsecond 


kW 


kilowatt 


kWh 


kilowatthour 


kWh/°C-kg 


kilowatthour per degree Celsius 


L 


liter 


lb 


pound 


lb/ft 


pound per foot 


lb/Mft 


pound per thousand feet 


lb/mi 


pound per mile 


lb/yd 


pound per yard 


L/h 


liter per hour 


m 


meter 


m 2 


square meter 


mA 


milliampere 


mA 


microampere 


m f 


microfarad 


mH 


millihenry 


pH/ft 


microhenry per foot 


min 


minute 


jtin 


microinch 


mJ 


millijoule 


mL/h 


milliliter per hour 


mm 


millmeter 


mm 2 


square millimeter 


MQ 


megohm 


mU 


milliohm 


nil 


microhm 


MPa 


megapascal 


mi/h 


mile per hour 


ms 


millisecond 


m/s 


meter per second 


m/s 2 


meter per second squared 


MS 


microsecond 


MV 


megavolt 


mV 


millivolt 


*»V 


microvolt 


MVA 


megavoltampere 


MV//1S 


megavolt per microsecond 


N 


newton 


N-m 


newton meter 


n 


ohm 


n/°c 


ohm per degree Celsius 


Q-cm 


ohm-centimeter 


fl/cm 2 


ohm per square centimeter 


Q-cmil 


ohm-circular-mil 


fi-cmil-ft 


ohm-circular-mil-foot 


O-cmil/ft 


ohm-circular-mil per foot 


a/cmil-ft 


ohm per circular-mil-foot 


fi-ft 


ohm-foot 


fi-in 


ohm-inch 


Q-m 


ohm-meter 


O/Mft 


ohm per thousand feet 


fi/mi 


ohm per mile 


Q/V 


ohm per volt 


oz/ft 3 


ounce per cubic foot 


pC 


picocoulomb 


pF 


picofarad 


psi 


pound per square inch 


psia 


pound per square inch, absolute 


psig 


pound per square inch, gauge 


rad 


radian 


rad/s 


radian per second 



417 



r/min 


revolution per minute 


r/s 


revolution per second 


S 


Siemens 


s 


second 


S/m 


Siemens per meter 


T 


tesla 


ton/ft 2 


ton per square foot 


V 


volt 


VA 


voltampere 


VIA 


volt per ampere 


Vac 


volt, alternating current 


var 


voltampere reactive 


V/cm 


volt per centimeter 


Vdc 


volt, direct current 


V/m 


volt per meter 



V/Mft 


volt per thousand feet 


V/mi 


volt per mile 


V/jis 


volt per microsecond 


Vs 


volt-second 


W 


watt 


W/A 


watt per ampere 


Wb 


weber 


WbA 


weberampere 


Wb/A 


weber per ampere 


Wb/m 2 


weber per square meter 


Wh 


watthour 


W/(m-°C) 


watt per meter degree Celsius 


wt% 


weight percent 


yd 3 


cubic yard 


yr 


year 



OTHER ABBREVIATIONS AND ACRONYMS 



ac alternating current 

ACSR aluminum conductor steel reinforced 

AMSW ammeter switch 

AWG American Wire Gauge 

BIL basic impulse insulation level 

BS breaking strength 

cemf counterelectromotive force 

CFR Code of Federal Regulations (U.S.) 

ckt bkr circuit breaker 

CRT cathode-ray tube 

CSP chlorosulfonated polyethylene 

CT current transformer 

dc direct current 

DS disconnect switch 

emf electromotive force 

EPR ethylene propylene rubber 

FA forced air 

FET field-effect transistor 

FOA forced oil and air 

GCR ground-check relay 

GCS ground-check system 

GND ground 

GTR ground-trip relay 

h.s. high-strength (guy grade) 

IC integrated circuit 

ICEA Insulated Cable Engineers Association 

IEEE Institute of Electrical and Electronics Engineers 

LCD liquid-crystal display 

LED light-emitting diode 

MCC motor control center 

MCM thousand circular mils (wire gauge) 

MESA Mine Enforcement and Safety Administration (U.S.) 

MESG maximum experimental safe gap 

m-g motor-generator 

MOS metal oxide semiconductor 

MOV metal oxide varistor 

MSHA Mine Safety and Health Administration (U.S.) 

NBR nitrile butadiene rubber 

NC normally closed 



NEC National Electrical Code 

NEMA National Electrical Manufacturers Association 

NESC National Electrical Safety Code 

NO normally open 

NTIS National Technical Information Service, U.S. 
Department of Commerce 

OA over air (self-cooled) 

OCB oil circuit breaker 

OL overload relay, overhead line 

OTR overtemperature relay 

PCB polychlorinated biphenyl 

pf power factor 

PIV peak inverse voltage 

PM preventive maintenance 

PT potential transformer 

pu per unit 

PVC polyvinyl chloride 

RC remote control 

RFI radio frequency interference 

RM rotating machinery 

rms root-mean-square 

SA surge arrester 

SBR styrene butadiene rubber 

SCR silicon-controlled rectifier 

SEL sensitive earth-leakage (system) 

SF safety factor 

SI International System of Units 

s.m. Siemens-Martin (guy grade) 

TBS two-breaker skid 

TDR time-domain reflectometer 

tpdt three-pole double throw (switch) 

UV undervoltage 

UVR undervoltage relay 

VCB vacuum circuit breaker 

VOM volt-ohm-milliammeter 

VR voltage regulator 

VTVM vacuum-tube voltmeter 

WVDC working volts, direct current 

XLP crosslinked polyethylene 



418 



ELECTRICAL AND ELECTRONICS SYMBOLS 



A 
a 



D 

d 



F 
f 



GC 
H 

h 

I 
I 



J 
K 



k 

L 

I 

LF 



anode 
area 

turns ratio of transformer (chapter 3) 
unit vector eJ 120 (chapter 4) 
radius of rod, spacing between electrodes 
(chapter 7) 

susceptance, magnetic field flux density 
area (chapter 7) 
base (chapter 8) 

capacitance 
collector (chapter 5) 
capacity (chapter 8) 
controlled variable (chapter 14) 

electric flux density (chapter 5) 
diameter, in feet or meters (chapter 8) 
diode (chapter 9) 
conductor diameter; coil diameter, in inches 

(chapter 2) 
distance (chapter 11) 

electric field strength 
emitter (chapter 5) 
potential (chapter 7) 
difference or error (chapter 14) 
dielectric constant (chapter 17) 

magnetomotive force 
factor of safety (chapter 13) 
force, frequency (chapter 6) 
safe bending stress (chapter 8) 

conductance 

cable code: grounding conductor 
gate, galvanometer (chapter 5) 
ground-fault device (chapter 9) 
cable code: ground-check cable 

magnetic field strength 

feedback element (chapter 14) 

thickness 

total operation time for motor (chapter 6) 

current 

current phasor, total circuit current 

instantaneous current 

current density at electrode surface (chapter 7) 

current through circuit breaker, system current 

(chapter 11) 
conjugate of complex current (chapter 3) 

imaginary operator, V^T (chapter 2) 

Boltzmann constant (1.38 x 10- 23 J/K) 

(chapter 5) 
torque constant, proportionally constant 

(chapter 6) 
reflectance factor (chapter 7) 
winding series capacitance (chapter 11) 
cathode (chapter 14) 
relay (chapter 15) 
coefficient of coupling (chapter 3) 

inductance 

total length, distance, in feet or meters 

(chapters 7-8) 
length, in inches (chapter 2) 
path of ionization (chapter 17) 
actual average power consumed divided by rated 

average power (chapter 8) 



M 



MP 

MPF 

MP-GC 



N 



Q 
q 

R 



U 

y 
v 

v 
W 



X 
x 



mechanical moment 
mutual inductance (chapter 3) 
machine contractor (chapter 9) 
voltage-sensing relay (chapter 12) 
cable code: mine power 
cable code: mine power feeder 
cable code: mine power feeder with ground-check 
conductor 

north 

number 

turns of the coil (chapter 2) 

number 

semiconductor with excess negative charge 

(chapter 5) 
armature speed (chapter 6) 

power, permeance 
instantaneous power (chapter 2) 
semiconductor with excess positive charge 

(chapter 5) 
number of magnetic poles presented by stator 

(chapter 6) 

charge, reactive power 
electrical charge, stored electrical charge in 
capacitor 

resistance, reluctance 

reference variable (chapter 14) 

radial distance or moment arm (chapter 6) 

equivalent radius (chapter 7) 

apparent power 

south 

distance (chapter 8) 

transformer capacity (chapter 11) 

displacement (chapter 2) 

motor slip (chapter 6) 



s 


complex power 


SH 


cable code: shielded 


SH-C 


cable code; see tables 8.2, 8.3 


SHC-GC 


cable code; see tables 8.2, 8.3 


SH-D 


cable code; see tables 8.2, 8.3 


SHD-GC, 


cable codes; see tables 8.2, 8.3 


SHD+GC 




SS 


substrate 


T 


circuit configuration (chapter 2) 




junction temperature (chapter 5) 




torque (chapter 6) 




time (chapter 7) 




tensile strength, tension (chapter 8) 


t 


temperature 




time 




taper (chapter 8) 


Ta 


ambient temperature (chapter 5) 



velocity of propagation (chapter 11) 

voltage, electromotive force, potential difference 
voltage phasor 
instantaneous voltage 

energy, work 

weight 

wind (chapter 8) 

cable code; see tables 8.2, 8.3 

work done (chapter 2) 

reactance 

distance (chapter 11) 



419 



Y 


admittance 


e 


y 


wye circuit configuration 




z 


impedance 




|Z| 


magnitude of impedance 




z 


burial depth 




a 


temperature coefficient (chapter 2) 


X 

\ 




no-load position, angular position (chapter 6) 


n 




firing angle (chapter 14) 


Mi 




ionizing coefficient of ions (chapter 17) 


P 


|3 


a divided by (1 - a) 


a 


7 


conductivity 




A 


delta circuit configuration 


$ 


5 


soil density (chapter 7) 




e 


permittivity 


</> 


>7 


efficiency (chapter 3) 





phase angle 

maximum allowable soil temperature rise 

(chapter 7) 
power-factor angle (chapters 4-5) 
protective angle (chapter 11) 
thermal resistance 
relative permittivity 
soil thermal conductivity (chapter 7) 
permeability (absolute) 
relative permeability 
resistivity 
conductivity 

soil specific heat (chapter 7) 
potential difference, magnetic flux 
flux direction (chapter 6) 
phase angle (chapter 2) 
flux (chapter 15) 
electric flux 
angular frequency 



420 



INDEX 
-A- 



Page 

Abnormal transient 282, 355 

Accuracy 75, 118-120, 174, 208, 246-247, 

270, 272-274, 331, 361, 362 

Active power 63, 137, 319 

Admittance 55, 64, 72, 268 

Air circuit breaker 87, 227 

Alternating current 2, 45, 48, 50, 129, 244, 

246-247, 301, 347, 351 
Alternating current mine power centers . . 280, 302-307, 310, 313, 

315-317, 319-320, 
325, 331-345 

couplers 9, 13, 15-16, 191-195, 206, 208, 222, 224, 252, 

254, 303-304, 317, 319, 360, 377-381 

protective circuitry 13, 75, 98, 103, 159, 165, 

179-180, 186, 224, 247 
Alternating current reclosing breaker (see Recloser) 
Alternating current time-overcurrent relay . . 90, 166, 243-245, 247, 

249-250, 256, 270- 

271, 328-330, 334, 

358, 360-362, 

364-366 

Aluminum 122, 130, 174, 178, 184, 194, 196, 203, 

211, 216-217, 242, 244, 312, 339 

bus 5, 7, 9, 10-11, 13, 108-109, 201, 224, 236, 255, 262, 

264, 266, 268, 306, 308, 310, 312, 314, 322, 344 

cables 184 

connectors 5, 191-192, 208-210, 228-231, 313, 376-378 

Faraday shields 298, 311, 379 

transformer windings ... 68, 70-71, 74, 86, 95, 109, 180, 290, 

308-311, 321-322, 341-383 

Ambient temperatures, correction factor 195-196, 271, 399 

American Wire Gage (AWG) 161, 184 



Page 

Ammeters 51, 72, 89-90, 115-117, 119-120, 123-124, 127- 

128, 273-274, 309, 317, 329, 344, 380, 398 

Ampacity 195-199, 201, 255, 271, 276, 311, 313-314, 331 

Ampere 20, 86, 93, 95-96, 103, 119, 154, 160-162, 171, 

196-197, 229, 235, 240, 260, 262, 271 

Ampere-turn 245, 272-273, 314, 331 

Amplifiers 22, 109-114, 127, 353, 362 

Apparent power 59-60, 62-63, 66, 68, 74, 81, 

83, 93, 103, 118, 246, 319 

Approval 186-187, 194, 207, 303, 382-383, 391-393 

Arc . . 3, 97-98, 125, 150, 159-161, 166, 181, 192, 211, 215, 217, 224- 

229, 232-234, 237-238, 255, 261, 275, 281-282, 285-287, 

290, 292, 299-300, 306, 324-325, 335-336, 338, 360, 

362, 367, 370, 376, 380-381, 384, 389-391, 393-395 

Arc, incendive 362-384 

Arc chutes 227-229 

Arc fault 98, 261, 384, 390 

Arc interruption 225, 227, 229, 232-234, 335 

Arc quenching 98, 227, 229 

Arc tracking 192, 390 

Arcing, intermachines 166, 255, 384 

Armature 1, 77, 130-132, 136, 146-154, 

229, 241-242, 324, 348, 350 

Armature reaction 147-148 

Armature windings 130-132, 136, 146, 148-149 

Arresters, surge 22, 292-296, 298-299, 301, 306-307, 310, 

327, 332, 334, 337, 343, 396 

Asymmetrical current 229, 262-263, 268, 313, 329, 344 

Atoms 20, 104 

Autotransformers 74, 114, 141, 159 

Availability 177, 217, 220, 225, 233, 329, 332, 396-397 

Avalanche diodes 298 



-B- 



Backup relaying 254-256, 315, 360 

Base line 9-11, 93, 219 

Basic impulse insulation level (BIL) 290, 307, 336 

Basic power circuit 76, 84 

Battery 2, 42, 87, 117, 124, 192, 198, 330-331, 

350, 359, 366-381, 391-392, 400 

Battery boxes 374, 376, 392 

Battery charging 367, 371-373, 377, 379-381 

Battery tripping 331 

Belt-conveyor starters 319, 353-355, 360, 365 

Bias 104-106, 109-114, 199, 257, 346 

Bidirection thyristor control 348 

Bimetal 229-232 

Blowout coils 324-325 

Bolted fault 98, 102, 261, 263, 270, 275, 322 

Bonds 164-166, 185, 207-208, 210-211, 215, 

229, 252, 290, 339, 345, 367, 379 

rails 216 

welded 216 

Borehole cable installation 203 

Boric acid fuse 238-239, 276 



Braking, dynamic 150-158 

Branch 8, 35-39, 4M2, 45, 56-57, 65, 80, 

220, 247, 274, 326-327, 329 

Brazed connection 179 

Breakers (see Circuit breakers) 

Breakover voltage 144, 346 

Bridge circuits 34-37 

Kelvin double 123, 398 

Wheatstone 122 

Bridge rectifiers 106, 143, 321-322, 324, 330, 348, 363 

Broken-delta relaying 251 

Brushes (see Motor brushes) 

Bucket-wheel excavators 8 

Burden 8, 118-120, 246-247, 272-274, 314, 

328-329, 331, 361-362, 365, 396 

Bus 5, 9-12, 108, 201, 236, 255, 262, 264, 

266, 268, 306, 310, 312, 322, 344 

protection 7, 13, 314 

Bus bar 109, 224 

Bus bar connections 244 

Butt-wrap grounding 299 



421 



-c- 



Page 



Page 



Cable-fault location 98, 102, 159, 207, 244, 263- 

264, 266, 268, 328, 376 

Cable-reel locomotives 367 

Cables 1, 3, 14, 29, 81, 85, 87, 91, 101, 109, 162, 165-166, 

216-220, 228, 230, 271, 275-276, 278-279, 283- 

288, 297, 305, 307-310, 312, 319, 323, 329, 

332, 354, 362, 377-380, 382, 384, 387-388 

alternating current mine 2, 46, 48, 50, 129, 244, 

246-247, 301, 351, 366 

aluminum 184 

ampacity 195-198, 201, 271, 313 

battery 375 

borehole 13, 186, 191, 203 

capacitance 30, 268, 282, 285 

charger 198, 350, 367, 370-373, 375-381 

clamps, in mine shafts 193, 203-205 

copper 184, 213 

corona in 185, 192 

couplers 5, 191, 222, 254, 304, 317, 379 

direct current mine 30, 129, 131, 135, 158, 252, 

256, 320, 324, 350 

drag 183, 190, 205-207 

entrance 193, 391 

fault current 97-98, 102, 159-161, 171, 180, 218, 224-226, 

229, 236, 243, 246, 254-255, 257, 260-264, 

266-268, 270-277, 287, 292, 306, 310-315, 

322-323, 329, 336, 339, 363, 375 

feeder 9, 16, 76, 96, 103, 124, 154, 182, 184-185, 195, 

197, 204, 211, 213-214, 260, 270, 289, 320 

flameproof 395 

flat 187-188, 193, 197, 210 

high-voltage 13, 193 

installation 193-194, 202, 204, 296 

insulation 183, 185-187, 190, 193, 196, 199, 

205, 207, 209-211, 398, 401, 403 

interlocked armor 191, 203-204 

jackets 183, 185-187, 193-195, 199, 203, 

205-206, 208, 210, 393 

locating faults 207 

maintenance 202 

mine power feeder 191 

portable 2, 182, 184, 187, 193, 197, 201, 204, 

206, 209, 222, 290, 396, 398 

rating 159, 185, 195, 199 

reactance 265 

resistance 123, 164, 207 

round 187, 188, 190, 193, 197-198, 204 

shielded 98, 186, 190-191, 210 

shuttle-car 183, 367 

size selection 200, 202 

splicing 183, 207, 210, 222, 258, 290 

stresses 183, 191 

symmetrical . . 82-83, 97-103, 117, 179, 186, 190, 225, 229, 232, 
235-237, 239, 246, 250, 261-263, 266, 268, 270, 
273, 298, 310, 313, 329, 331, 346, 354, 404 
trailing (see Trailing cables) 

types 182, 186, 187, 190, 204, 257 

voltage-drop criteria 195, 202 

Capacitance 21, 28-31, 33, 45^8, 50-51, 53, 56-57, 59, 61-64, 

81-82, 107, 115, 122, 160, 167, 170, 179, 186, 

207, 260, 268, 276, 280, 282-288, 290-291, 

295-298, 311, 319, 351, 359, 360, 

363, 371, 383-384, 399 



Capacitance switching 282-284, 286, 301, 332, 359 

Capacitive circuit 301-302, 384 

Capacitive reactance 54, 297 

Capacitive susceptance 47 

Capacitor motor 157 

Capacitor-trip device 329 

Capacitors 28, 30, 47-48, 54, 106, 111-113, 115, 124, 147, 

154, 156-157, 207, 237, 248, 253, 280, 
284, 292, 298, 328-331, 359, 399^00 

power-factor improvements 283, 332 

protection 296, 319 

rating 2% 

surge 87, 260, 268, 282, 295-297, 307, 332 

Capacity, interrupting 18, 224, 228-229, 232-234, 238, 268, 

305, 310, 313, 325, 329, 336 

Catalyst battery caps 375, 380 

Cathode ray oscilloscope 126 

Cathode spot 226-227, 233 

Cathodic protection 178, 181 

Certification 382, 389, 391 

Chain conveyor, longwall 12 

Characteristic impedance 284, 286, 288, 295, 307, 332 

Charge 20-21, 28-29, 47, 104-105, 115, 126, 167, 179, 

207, 225, 260, 280, 282, 284, 290, 295, 329- 
330, 346, 368-370, 372-374, 376-377, 380 

Charge cycle 369-370, 372-373, 380 

Chargers 350, 367, 370-373, 375, 377-381 

Charging, battery 367-368, 373, 377 

Charging current 179, 260, 268, 284, 296, 297, 370, 380 

Charging station 372-373, 375, 377, 380 

Chemical treatment, ground beds 177, 180 

Chopper motor control 349-350 

Chopping transient 285, 296 

Circuit breaker 7, 9-10, 13, 75, 86, 97-98, 179-180, 184, 224, 

239-240, 244, 247-248, 250-256, 259-260, 

263, 267-279, 282-283, 286, 290-292, 

298, 301, 304-307, 310, 312-319, 

322, 326-327, 329-336, 342, 

344, 353, 356, 360-367, 

377-378, 400, 402 

air 87, 227 

air magnetic 227-228, 232, 324 

alternating current 232 

contactor 90, 114, 150-151, 159, 257, 287, 324- 

325, 350, 359, 379-380, 393, 397 

direct current 232 

frame size 228-230, 232 

high-voltage 226-227, 232-233 

low-voltage 226-228, 231 

magnetic 227, 229-230 

medium-voltage 226-228, 231 

mine duty 229 

minimum oil 232-234 

molded case 226-232 

oil, dead tank 233, 329, 335 

oil, live tank 233 

power 226-228, 332 

ratings 227-228, 232-233 

high-voltage 232-233 

low-voltage 227 

medium-voltage 227 

molded case 228 

oil 232-233 



422 



Page 



Page 



power 232 

thermal-magnetic 229, 230-231 

vacuum 5, 232, 234 

Circuit reduction 31, 34, 36, 39, 41, 44, 56 

Circuits ... 2, 5-9, 11, 13, 19-50, 53-59, 61-73, 75-79, 82-86, 89, 91, 

93-95, 97-98, 100-117, 119-120, 122-125, 127-129, 136, 

144-146, 150-152, 154, 159-161, 163, 179-182, 200, 

215, 224-232, 234-242, 244, 246-263, 266-268, 

270-272, 274, 278, 280, 282-292, 294-296, 

298, 300-308, 310-319, 321-337, 339, 

346, 348-365, 367-368, 370, 376- 

380, 383-384, 389-390, 392, 

393-396, 398, 404 

alternating current 53, 57-58 

bridge 34-37, 123 

common-base 110-111 

common-collector 110, 112 

common-emitter 110-112 

comparator 363-364 

control 75, 141, 151, 154, 316, 324, 327, 

329-330, 331, 349, 354, 370 

direct current 56 

distribution 237-238, 260, 308, 334, 343 

integrated 114, 298, 364 

magnetic 274, 311, 324 

parallel 24-25, 44 

protection 184, 216 

rectifier 105-109, 117, 321, 346 

regulator 349 

series 22-25, 54, 76, 94 

solid-state 360, 362 

thyristor 113-114, 237, 287, 298, 323, 346-360, 370-372 

transistor 109-114, 298, 356-359 

triac 356-360 

tripping 184, 216, 255, 305, 310, 318, 329- 

330, 344, 356, 359, 367, 378 

Circular mil 22, 184 

Clapper relay 241-243, 356 

Clear 19, 73, 90, 161, 220-221, 224, 228, 232, 

234, 285-286, 306, 311, 329, 357 

Close-and-latch current 225, 229, 267, 268, 305 

Coal dust ignition 12, 17, 367, 382, 389, 391-395, 399 

Coal preparation plants 394-395 

Code, National Electrical (NEC) 195, 382, 395 

Code of Federal Regulations (CFR) 19, 182, 207, 222, 

259, 278, 395 

Coefficient of coupling 66, 69 

Coefficient of diffusion 374 

Coefficient of grounding 293 

Coils 29, 66, 69-70, 82, 115-120, 124, 133, 136, 138-139, 145, 

150, 152, 231, 237, 241, 247-249, 251-253, 256-257, 

271-272, 287, 290, 313-316, 323, 328, 330, 

336-337, 356-361, 372-373, 398, 404 

blowout 227, 324-325 

long 26-27 

polarity 22, 40, 65, 90, 98, 104-106, 109, 131-132, 161, 

242, 244, 246, 258, 280, 318, 377, 380, 390 

toroidal 28 

Coincident demand 103 

Commutating 

diodes 348 

fields 148 

poles 148 

Commutation 148, 322, 348 

Commutator 4, 131-132, 136, 147-148, 154, 350, 404 

Compensating winding 148 



Complex algebra 48-51, 53, 55 

Complex power 59-62, 68, 73-74, 81 

Complex quantity 52-55, 59 

Components, symmetrical 98-103, 179, 250, 261- 

262, 266, 268, 404 

Compound generator 132 

Compound motor 148-149, 151 

Computers, use of 114-115, 125-126, 200, 202, 260, 263, 

268, 278, 2%, 364, 398, 404 

Concentration 9, 13, 15, 29, 104-105, 127, 133, 175, 287, 

300, 344, 373-375, 382, 389, 393-394 

Conductance 22, 25, 32-33, 55, 64, 67, 292 

Conductivity 104, 165, 171, 174-175, 177-178, 184, 

201, 211-212, 251, 311-312, 399 

Conductors . . 2, 5, 8, 13-14, 20-21, 26-29, 65, 68, 73-75, 77, 79, 85- 

86, 91, 97-98, 100-104, 113, 118, 120, 124, 127, 129, 

132, 133-134, 136, 137, 141-142, 146-149, 154, 156, 

160-164, 166, 167, 171, 172, 174, 178, 180-183, 

192, 199, 200, 208, 211, 213, 216, 224, 227, 

229, 231, 233, 236-237, 244-250, 252-253, 

256, 259, 261, 273, 276, 280-281, 287- 

288, 290, 292, 299, 304-305, 307, 310- 

312, 314-318, 324, 328-331, 334, 

338-342, 354, 375, 377-380, 383- 

384, 388, 390-393, 400, 404 

aluminum 184, 194, 196, 312, 313 

ampacity 251, 275-276 

cable 9, 81, 165, 184-185, 187, 194-195, 210, 217, 

220, 257-258, 262, 268, 275, 329, 332, 401 

capacitance 167, 170, 268 

grounding 159, 165-166, 186-187, 191, 193, 

218, 255, 257-258, 319, 323 

impedance 101, 179 

inductance 170, 176, 273, 294 

insulation 71, 130, 138-139, 183, 185, 186-187, 194, 

202, 204, 206-207, 209-210, 271, 344 

reactance 139 

resistance 22, 31, 70, 123, 139-140, 159, 195 

shields 186-187, 190, 298, 337 

size 3, 184, 195-197, 199, 202, 217, 230, 264, 275 

spacing, overhead lines 217 

span, overhead lines 217-221 

stranded 185, 201 

strength 184, 202-203, 209 

Conduit 219, 297, 388, 393 

Conjugate 49, 60, 97 

Connected load 5, 84, 100-101, 103, 151, 270, 285, 308, 334 

Connections 3, 9, 78-79, 82, 84, 86, 92, 95, 116, 118-120, 124, 

128, 132, 136, 148-149, 151, 157, 159-160, 181, 

187, 192-193, 207-209, 211, 224, 231, 244, 246- 

249, 250-251, 261, 272, 289, 294, 2%, 299, 

307-308, 313-314, 316, 328, 332, 337, 339, 

341, 343, 356, 379-380, 387, 399, 401-402 

Connectors 5, 191-192, 208-209, 228, 231, 313, 376-377, 391 

Contact ... 98, 129, 132, 141, 143-145, 150-152, 159-162, 164, 166, 

169, 183, 186-187, 190-194, 204-205, 209-210, 215- 

216, 218-222, 225, 226-229, 231, 233-235, 237, 

240-243, 244-246, 248, 251-254, 256-259, 263, 

268, 270-271, 278, 280, 282, 284-287, 295, 

298-299, 302, 304-306, 310, 312, 314, 

317-319, 323-325, 328-330, 336, 339, 

353, 356-360, 363, 375, 377-379, 

381-385, 389-390, 392-393, 398 



423 



Page 



Page 



bounce 286, 360 

breaker 248, 254, 259, 285-286, 323-324, 328-329 

relay 240, 242, 253, 314, 330, 358-359 

switch 253, 305 

whiskers 286 

Contactor 90, 114, 150-151, 159, 257, 287, 324- 

325, 350, 359, 379-380, 393, 397 

Continuity of service 6, 276 

Continuous of current . . . 195-197, 225-226, 228-234, 237-238, 245, 

260, 274, 276, 305-306, 308, 312-314, 

325, 329, 331, 335-337, 355-356, 362 

Continuous miner ... 3, 12-13, 15, 81, 103, 153-154, 190, 198-199, 

201, 209, 308, 313-314, 319, 350-351, 391 

Continuous rating 153, 228, 231, 236, 239 

Control systems 4, 349, 352-354, 365 

Control transformer 306, 315-316, 324, 326, 329, 343 

Control wiring 224, 226, 248 

Controllers (see Starters) 

Convenience outlets 247, 316 

Conventional mining 3, 11-13, 182, 367 

Conveyors 3, 8, 12-13, 16, 113, 140-141, 143, 153, 182, 308, 

342, 350-354, 356, 360, 367, 392, 394-395, 397 

Cooling 7, 106, 134, 155, 183, 191, 197, 227-228, 233, 238, 

307, 310-311, 334, 350, 354, 384-385, 387 

transformers 307, 311, 334 

Coordination, protective relaying 254, 336 

Coordination-curve plots 276-278 

Copper 22, 70, 136, 139, 161, 177-179, 184-186, 190, 193, 

195-196, 200-203, 211-213, 215-216, 226, 
231, 236, 307, 311-313, 339, 390, 398 

cables 184, 213 

Faraday shields 186, 201, 311 

Core loss 69-70, 72-73, 137, 310 

Cores 27, 65-66, 68-70, 72-75, 82, 109, 119, 124, 127, 133, 

136-140, 149, 177, 208, 246, 250, 258-259, 272, 
285, 310-311, 324, 361, 370-371, 379, 402 

rotor 130 

stator 132, 137-139 

transformer 75, 87 

Corona 185, 192 

Corrosion 180, 184, 192, 194, 203, 375-376, 380 

conductor 178, 339 

electrode 177-178, 181 

explosion-proof enclosures 392 

ground-bed conductors 177, 180 

Coulomb 20 

Counterelectromotive force (cemf) 149 

Counterpoise 299, 301 

Couplers, cable (see Cable couplers) 

Coupling 9, 64-67, 69, 119, 160, 163, 192, 194, 

291, 317-318, 338, 360, 398 

magnetic 27, 65-67, 105, 145 

motor 145 

Crest voltage 47, 51, 285-286, 292-293, 295-296, 306 

Cross bonding 166, 216 

Cross field, motor 156-157 

Crosslinked polyethylene insulation 185 

Crowbars 298 

Cumulative compound motor 151 



Current . . 2-3, 18, 20-48, 50-57, 59-76, 78-95, 97-127, 129, 131-133, 

135-142, 145-157, 159, 160-167, 170-176, 178, 179-181, 

183-186, 190-192, 194-199, 201-202, 206-207, 211, 

215-216, 218, 220-222, 224-232, 234-268, 270- 

290, 292-301, 305-309, 310-326, 328-329, 

331-332, 335-341, 343-344, 346, 348- 

365, 367-373, 375-380, 384, 390, 

392, 394-395, 398^00, 403 

asymmetrical 229, 263, 267-268, 329, 344 

balance 25, 79, 249, 261 

carrying capacities 192, 196, 211, 216, 222 

bus 322 

charging 179, 260, 268, 284, 296-297, 370, 380 

chopping 284 

continuous 195-197, 225-226, 228-234, 237-238, 245, 260, 

274, 276, 305-306, 308, 312-314, 325, 
329, 331, 335-337, 355-356, 362 

definition 20 

density 26, 119, 129, 133, 137, 140, 148, 167, 171 

earth 166, 323 

eddy 70, 136-137, 311 

exciting 69-73, 94-95, 274, 295, 308 

fault 97-98, 102, 159-161, 171, 180, 218, 224-226, 229, 236, 

243, 246, 254-255, 257, 260-264, 266-268, 270-273, 

275-277, 287, 292, 306, 310-313, 315, 322- 

323, 329, 331, 336, 339, 363, 375 

inrush, motor 277, 283 

inrush, transformer 150, 244, 270, 276, 285- 

286, 295-296, 306, 337 

interrupting 225, 229, 232-234, 238, 268, 305, 

310, 313, 325, 329, 336 

leakage 190, 206-207, 356, 360, 375, 399 

line 79-80, 83-85, 92-93, 97, 100-102, 118, 120, 

137, 141, 146, 197, 225, 245, 249-250, 
254, 260-261, 263, 346, 353-354, 362 

locked-rotor 141 

loop 36-38, 57, 129 

magnetizing 69-70, 72, 284-285 

momentary 225, 360 

motor-starting 125, 143, 145, 151, 157, 200, 236, 238- 

239, 277, 283, 298, 318, 363-365 

overload 274, 275 

phase 79-80, 83-84, 92, 101, 108, 120, 250 

short-circuit . . 44, 73, 202, 225, 229, 232, 236-237, 260-264, 270, 
275, 277, 310, 313, 321, 328, 331, 336, 344 

symmetrical three-phase 263, 313 

transient 283 

unbalanced 102, 247 

Current-balancing transformers 109 

Current-limiting capacity 18, 87, 224 

Current-limiting fuses 235-237, 239, 247, 276, 306, 338 

Current-limiting resistors 165 

Current transformers 83, 108 

accuracy 118, 246, 272-273 

burden 118, 246, 272-274, 328-329 

errors 272-275 

model 272 

ratios 244-246, 272-274 

relaying 89, 118, 244-246, 249, 272-274, 328-329 

saturation 272-273 

Cutting machine 12-13, 183, 187, 198 

Cylindrical-rotor motor 144-145 



424 



-D- 



Page 



Page 



Damper winding 145 

Damping 242-243, 283, 285 

D'Arsonval meter 115-117, 119-120, 124-125 

Data ... 19, 86, 126, 180-181, 197-199, 202, 211, 222, 266, 273-274, 
277, 331, 364-365, 381, 387, 393, 395, 397-399, 404 

cable 213, 271 

Dead lock 14 

Dead front 312, 316-317 

Decrement factor 267 

Deep-bar rotor (see Rotor) 

Definite-time relay 151 

Delta connections 35, 79-80, 82, 84-85, 100, 108, 

138, 141, 258, 296, 308, 321 

Delta-delta transformers 83, 92, 261, 308, 343-344 

Delta-wye transformations 34, 83-84, 92 

Delta-wye transformers 90, 92-93, 261, 308-309, 344 

Demand 24, 17, 68, 103, 122, 153-155, 197, 199, 

201, 255, 271, 275-276, 302-303, 308, 
312, 314, 331, 335, 345, 350, 364 

Demand factor 103, 308, 335 

Demand meter 90, 122, 127, 345 

Diagrams, one-line 86, 91-92, 95-96, 163, 201, 255, 262, 264, 

276-277, 302-303, 333-334, 336, 344 

Dielectric 28, 186, 191-192, 210, 221, 228, 234, 280, 285- 

286, 290, 292, 295, 300, 398400, 403 

Different-current relaying (direct current) 36, 323-324, 336, 

337, 367 

Differential compound motor 151 

Differential protective relaying 90 

Diode 3, 22, 90, 104-105, 108-112, 114, 125, 165, 181, 

237, 257-258, 275, 287, 298, 322-323, 329, 
346, 348-349, 354, 363, 370 

Diode grounding 164, 187, 194, 256-258, 275, 323 

Direct current cables 202, 256 

Direct current circuits 21-22, 28, 30-31, 45, 50, 53, 55-56, 58, 

107, 117, 166, 227, 232, 242, 247, 

253, 258-259, 286, 298, 321- 

322, 324, 378 

Direct current circuit breaker 90, 228 

Direct current generators 2, 90, 133, 147-148, 150, 152 

commutator 131-132 

excitation 132 

Direct current ground-fault relaying 259, 323 

Direct current motors 3, 147, 156, 198, 287, 351, 380, 401 

chopper-driven 349-350 

compound 148-149, 151, 153, 348 

separately excited 4, 132, 148, 152 

series 2, 4, 135, 148, 151-154, 320, 348 

shunt 148-151 

Direct current offset current 166 



Direct current overcurrent relay 90, 247, 256, 361 

Direct relaying 246-250, 255 

Directional relay 88, 90, 240, 242, 244 

Discharge, partial 185-186, 192, 393, 403 

Disconnect switch 5, 9, 13, 220, 226, 237, 254, 256, 

305-306, 326-327, 336, 338, 393 

Discontinuity 207, 283-289 

Dissipation factor 399 

Distribution . . . 2-11, 13-15, 17-19, 22, 25, 30-31, 35, 55, 75, 78-79, 

82, 96, 103, 107, 124, 128, 131, 147, 153-154, 158- 

159, 161-162, 166, 174, 179, 181-183, 185-186, 

190-191, 197, 199-202, 211, 216-219, 222, 

224, 228, 234, 237-238, 244, 254-255, 

259-260, 262-263, 266, 272, 277-281, 

284, 286-287, 291-299, 301, 303, 

304-311, 316, 319, 321, 325-326, 

329, 331-338, 340, 342-345, 

350, 361-362, 365, 370, 

381-382, 391, 396, 398 

Distribution system ... 3, 7, 13, 124, 147, 153, 159, 166, 182, 190- 

191, 200-201, 211, 234, 272, 278-279, 281, 

292, 295-297, 305, 308, 319, 325, 329, 

331-332, 336, 344, 365, 370 

direct current 2, 164-165, 187-188, 226-227, 229, 232, 

247, 275, 286, 287, 298, 322, 350 

open pit mine 8, 11, 19, 185, 216, 219, 255, 301, 345 

preparation plant 8, 17-19, 129, 334, 340, 342, 

382, 384, 394-395, 397 

primary-selective 9 

radial 5-6, 10, 326, 333-334 

secondary-selective 6, 334 

strip mine 4-5, 9-11, 185, 216, 219-220, 232 

Distribution transformers 5, 7, 13, 75, 262, 297, 303, 309, 398 

Double-cage rotor 140-141 

Double-ended substation 7, 334 

Double switchhouse 5, 9, 326-327, 330 

Draglines 4-5, 8, 166, 182-183, 190 

Driving potential 30, 102 

Drum controller 150 

Dry-type transformer 293, 301, 337, 379 

Dual-element fuse 235-236, 239, 276, 312, 393 

Duality 64 

Dust-cap, receptacle 304-305 

Dust-ignition-proof enclosures 134, 382-383, 394 

Dust-tight enclosures 394 

Duty-cycle operation 153-156, 197-199, 314 

Dv/dt protection 323 

Dynamic breaking 150-152 

Dynamometer 115, 117-120, 122 



-E- 



Earth resistance 167-168, 174, 178, 180-181 

Earth resistivity 178, 180-181, 301, 341 

Eddy currents 70, 136-137, 311 

Efficiency 3, 12, 59, 71, 73, 77, 81, 84, 106, 137, 139- 

141, 154, 158, 186, 197-198, 232, 239, 

284, 292, 307-308, 313, 319, 329 

motors 138, 270 



transformers 146 

Electric charge 20, 28 

Electric field 28-29, 104-105, 167, 181, 186, 

225, 281, 286, 290, 394 

Electrochemical cell 178, 367 

Electrocution 208, 217, 219, 221, 317, 377-378 

Electrodes, grounding 172, 177, 181, 299, 301 



425 



Page 



Page 



Electromagnetic attraction 240-243, 356-358 

Electromagnetic induction 240, 242, 244, 356, 358 

Electromagnetic torque 133 

Electromechanical devices 240, 362 

Electromotive force (emf) 130, 290 

Electron theory 20 

Electrostatic force 20 

Electrostatic shielding 311 

Emergency-stop switch 254 

Enclosure 130, 159, 178, 220, 232, 236, 238, 240, 

302-305, 320, 326, 329, 332, 378-380 

battery 373-374, 377 

dust-tight 389, 394 

explosion-proof 382-387, 389-392, 395 

joints 385 

metal-clad 303, 326, 336 

motors 134 

Endosmosis 171 

Equations 22-23, 25-26, 28, 31, 34, 39, 42, 48-55, 59-66, 69, 

74-80, 84-85, 92-96, 99-101, 105-107, 110-111, 

120, 123, 137, 146-149, 155-156, 162, 170- 

171, 198, 263, 271, 282-283, 285, 288- 

289, 291, 295-297, 313, 368, 373 

circuit 30-36 

loop 36-40, 43, 57, 64, 67, 102 



node 38, 40, 43 

Equipment 1-19, 62-65, 75-76, 81-82, 85, 93-98, 103-108, 114- 

117, 120-125, 129, 135, 147, 153-154, 159-166, 

174, 178-187, 190-191, 194, 197-201, 204-206, 

218-222, 224, 228-236, 239-241, 246-248, 251, 

263-264, 265, 268-270, 274-281, 287- 

313, 316-328, 331-340, 343-345, 350- 

354, 360-362, 367, 375, 377-379, 

382-383, 387, 389-398, 401-404 

mine power 4 

mining 2 

Equivalence 25, 31, 33-34, 42, 80 

Equivalent circuits 40, 57, 68, 70-71, 73, 76, 85, 94-95, 113- 

114, 260-261, 266, 284-285, 291 

(EPR) insulation 185 

Error, current transformer (CT) 272-275 

Ethylene propylene rubber 185, 222 

Euler's theorem 49, 52 

Excavators, surface 129, 147, 152, 252, 278, 350 

Excitation 72, 90, 132, 143-146, 150-151, 260-261, 272, 348 

Exciter 90, 143, 147, 152 

Explosion proof 134, 191-192, 315, 382-392, 394-395 

Explosion-proof enclosures 382-387, 389-392, 395 

Extended-time rating 179, 312 

Extinction, corona 185, 192 



-F- 



Factors 3, 19, 52, 63-64, 81, 83-85, 90-91, 99, 105, 107, 110, 

117-119, 123, 127-128, 146-147, 153-154, 161, 166- 

167, 169-170, 174, 180, 184-185, 191, 195, 199- 

200, 201-203, 206, 208, 217, 219, 221, 225, 

234, 246, 261-264, 268, 270-272, 275, 278, 

280, 290, 293, 308, 311, 313, 319, 328, 

331, 335, 339, 342, 357, 363, 365, 367, 

372-374, 377, 380, 383, 386-387, 389- 

390, 392-393, 396, 398, 400, 403 

ambient temperature 196 

decrement 267 

demand 103 

dissipation 399 

diversity 103 

elevation 155 

load 61-62, 80, 103, 138-140, 197-198 

polarization 402 

power 59, 62, 114, 120, 138, 201, 399 

reflection 139, 295 

service 134 

Fall-of-potential measurements 172, 179 

Farad 295 

Faraday shield 298, 311, 379 

Faraday's law 129, 290 

Fault 88, 97-98, 101-102, 119, 159-163, 165-166, 171, 

179-181, 184, 186, 192, 207, 218, 220, 222, 

224-226, 229, 235-236, 243-244, 246-247, 

248-251, 253-255, 257-268, 270-273, 

275-278, 280, 287, 290, 292, 293, 

297-298, 301-302, 306, 308, 310- 

319, 322-325, 328-329, 331, 

333-334, 336-337, 339, 344, 

354, 360-364, 367, 375- 

380, 383-384, 390-391 



Fault, battery 375-377 

Fault, cable, locating 207 

Fault calculations 101, 260-263, 265, 267, 268, 383 

Fault current 97-98, 102, 159-161, 171, 180, 218, 224-226, 

229, 236, 243, 246, 254-255, 257, 260-264, 

266-268, 270-273, 275-277, 287, 292, 

306, 310-313, 315, 322-323, 329, 

331, 336, 339, 363, 375 

Fault-current sources 260 

Fault-point impedance 268 

Fault-through stress 224 

Feed-through receptacle 305, 326-327 

Feedback control 349-350, 353 

Feeders (electrical) . . 5-7, 9-10, 13, 15-16, 72, 76, 81, 96, 103, 124, 

154, 182-185, 187, 190-191, 195, 197-198, 

201, 204, 211, 213-215, 260, 270, 

287, 289, 320-321, 323, 333- 

334, 342, 345, 366-367 

Fencing 334-338 

Ferroresonance 287 

Ferroresonant transformers 371-372, 381 

Field-effect transistor (FET) 112 

Field emission 226, 234 

Field excitation 150, 348 

synchronous motors 4, 89, 135, 137, 143-148, 

152, 260-261, 268, 297 

Field strength, electric 29, 286 

Field windings 130-132, 143-146, 148-149, 151, 153 

Filter, rectifier 351 

Flameproof enclosure 383, 390, 395 

Flat-compounded motor 148-149, 151 

Flux, magnetic 26-27, 66, 69, 77, 109, 129, 

131-133, 148-149, 151, 280 



426 



Page 



Page 



Flux linkage 66, 69 

Forced response 48 

Forward bias 104-106, 109-111, 257, 346 

Forward blocking 346 

Frame, motor 133, 148 

Frame size, circuit breaker 228-230, 232 

Full-load currents 75, 140, 197-199, 247, 270-271, 273, 276, 

287, 309-310, 313-314, 322, 364 

Full-wave rectifier 106, 108, 117, 132, 364, 370 

Fuses 87, 97-98, 180, 224, 226, 244, 248, 254- 

257, 259, 266, 277-278, 292, 295, 312, 

316, 322, 329, 336, 338, 377, 380, 393 

boric acid 238-239 



class 235-237 

current-limiting 235-238, 284, 356 

dual-element 235-236 

expulsion 237 

high-voltage 235, 237-238, 276 

I 2 t 236 

low-voltage 235, 237-238, 263, 271, 276 

power 237 

ratings 235, 239, 260, 266, 270, 306, 322 

semiconductor 237 

transformer 247, 314, 333, 342 

Fusible element 98, 235-238, 306 



Galvanometers 123, 125 

Gap, explosion-proof enclosure 387 

Gaps, spark 292 

Gassing, battery 380 

Gate, thyristor 113-114, 346, 360 

General-purpose fuses 237, 239 

General-purpose motor 140 

Generator (see Direct current generators, synchronous machines) 

Ground 9, 13, 15, 18, 38, 79, 87-90, 106, 139, 153, 159-187, 

190-194, 201-204, 206, 208-221, 224-226, 233, 236- 

237, 247-261, 263, 268, 270-272, 275-276,278- 

281, 284, 286-287, 290-301, 304-309, 311- 

312, 314-319, 321, 323-325, 328-345, 

354, 361-363, 366-367, 375-380, 

383, 391-393, 400-402 

Ground bed, mesh or electrode . . 124, 160, 162-164, 166, 169-172, 

174-175, 177-181, 294, 299- 
300, 334, 338-341, 343 

Ground-bed measurement 172, 181 

Ground-check monitor 164 

Ground fault 159, 162, 165, 178-179, 181 

Ground-fault current 160-161, 180 

Ground insulation 120, 139 

Ground protective relay 90, 165, 179, 248-249 

Ground resistance 166, 169-170, 172-173 



Grounded conductor 159, 220, 257, 323, 393 

Grounded system 160, 251, 261, 271, 272, 276, 306 

effective 293 

noneffective 293 

resistance 163, 293, 319, 402 

solidly 293 

ungrounded 164, 284, 286 

Grounding 5, 18, 22, 82, 99, 119-120, 160, 162, 167, 170, 172, 

176-177, 181-182, 185, 208, 293-294, 299, 305- 
306, 325, 337, 339, 363, 380, 383, 391 

conductor 159, 161, 163-166, 171, 179-180, 184, 187, 190- 

191, 194, 206, 209-210, 217, 224, 247-253, 

255-256, 258, 268, 275-276, 281, 290, 304, 

311, 317-319, 323, 329, 330, 334, 

338, 340-342, 377-379, 392-393 

diode 257, 323 

instruments 180 

resistor 90, 161, 163-165, 179-180, 224, 247, 249-250, 253, 

255, 268, 272, 278, 287, 312, 314-315, 323, 
333-334, 340-341, 343-344, 354, 361 

system 1-2, 9, 13, 98, 159-161, 165-166, 169, 171- 

173, 178-179, 181, 251, 254, 271-272, 
281, 300-301, 332, 340-342, 345 

transformer 79, 179-180, 308-309, 321, 340, 343, 344 

Group motor control 17 

Guy wires 218 



-H- 



Half-cell 178 

Half-wave rectifier 106, 108, 329, 346-347 

Hall effect 126-127 

Harmonics 70, 286-287, 337, 346, 348, 354, 403-404 

Hazard reduction 383, 394-395 

Hazardous atmosphere 382-384 

Hazardous locations 382-383, 394-395 

Headlight lens, permissible 388 

Heat sinks 106-107, 110, 349-350, 355-356, 359, 386 

Heaters, strip 229, 329 

Heating of soils 175 

Henry 26-27 

Hertz 46, 124 

High field emission 225 



High voltage 3, 6, 8, 13, 15-17, 63, 71, 75, 120, 128, 135, 139, 

159-160, 165, 179, 181, 185-187, 190-195, 210- 

211, 217, 220-222, 226-227, 229, 232-235, 

237-239, 248, 250-251, 253-256, 260- 

263, 266, 267-268, 271-273, 276-277, 

280, 283, 285-287, 290, 293, 298, 

304-306, 308, 314, 317, 319, 326, 

328-329, 332-333, 336, 339- 

340, 342, 344, 362, 365, 

368, 390, 395, 402403 

High-voltage circuit breakers 13, 232, 239, 263, 268, 

273, 306, 329 

High-voltage couplers 191, 193-194, 304 

High-voltage distribution 217, 234, 250, 256, 266, 

272, 286, 298, 326, 365 



427 



Page 

Hipot 206-207 

History, mine electrical 2 

Hoists 129, 142, 152-153, 158, 218-219, 221, 278, 350 

Horsepower 2^t, 8, 84, 103, 134-135, 138, 140-142, 149, 

152-154, 197-198, 241, 260-261, 264-265, 
270, 308, 313, 335, 350-351, 354, 356 



Page 

Hot-spot temperatures 310 

Hybrid relay 356, 358, 360 

Hysteresis loss 70, 285, 311 



-I- 



I 2 t fuses 234 

Ideal sources 42-43, 57, 102 

Ideal transformers 66-71, 73-74, 91, 108, 363 

Idealization 29 

Imaginary numbers 48-49 

Imaginary power 59-60 

Impedance ... 20, 53-57, 61, 63, 64, 66-68, 71-74, 76, 78-80, 83-86, 

91-98, 101-102, 107, 109-113, 115, 118-120, 122- 

123, 125, 141, 156-157, 166, 179-180, 186, 200- 

202, 224, 241, 246, 252-253, 256, 258, 260- 

266, 268, 272-275, 276, 281, 284-289, 

294-296, 298-301, 307, 309-310, 312- 

313, 315, 317-319, 321-322, 332, 

335, 337, 342, 361-363 

Impedance angle 67, 85 

Impedance diagram 97, 266 

Impedance voltage 272, 309, 363 

Incendive arcing 362, 384 

Inductance ... 20, 26-31, 33, 45^18, 50-51, 53, 63-67, 69-73, 75, 118, 

122, 145, 156-157, 170, 174, 176, 230, 246, 280, 282- 

288, 290-291, 294-295, 324-325, 371, 383-384 

Induction-disk relays 136, 242-243, 249, 270-273, 278, 

328-329, 331, 361-362 

Induction motors 60, 62, 84, 89, 122, 135, 137, 140-141, 144- 

145, 152, 158, 242, 244, 260-261, 263, 

265, 267-268, 270-271, 297, 319, 

348, 351, 353-356, 363, 365 

single-phase 156-157 

three-phase 91, 136, 138-139 

wound-rotor 142-143, 283, 352, 397 

Inductive reactance 47, 54, 63, 139, 146 

Infrared 207, 40S404 

Instantaneous trips 229-230, 275-276, 312, 362 

Institute of Electrical and 
Electronics Engineers (IEEE) . . 18, 181, 199, 206, 222, 226, 245- 

246, 259, 261, 271, 273, 278- 

280, 301, 307-308, 325, 

334, 346, 365-366, 

379, 381, 382, 395 

Instruments 75, 89, 114-120, 122-128, 173- 

174, 176, 270, 399, 400, 403 



Insulated Cable Engineers Association (ICEA) [Insulated Power 

Cable Engineers Association (IPCEA)] 185-186, 194-201, 

204, 206-207 

Insulation 71, 119-120, 123, 135, 139, 155, 159, 179, 183, 

184-187, 190-196, 200-202, 204, 206-211, 

220, 246, 260, 271, 283, 290-295, 297, 

299, 303, 306-307, 310-312, 319, 325, 

328-329, 336-339, 345, 353-354, 

377-379, 382-383, 390, 391- 

392, 396, 398-404 

Insulation, cable 183, 185-186, 192, 195, 204, 

206-207, 209-210, 290, 404 

Insulation, conductor 139, 187, 195, 206, 208, 299 

Insulation, motors 135, 139 

Insulation, transformer 159 

Insulation class 135, 195, 246, 297, 310, 329, 337, 382 

Insulation coatings 378-379 

Insulation coordination 293, 337 

Insulation failure 138, 312 

Insulation shield 186-187, 190 

Insulation systems 135, 139, 186, 192, 290, 294, 306, 400, 401 

Insulator 20, 113, 166, 174, 179, 185-186, 192, 

194, 216, 218, 221, 233, 284, 312, 
332, 334, 344, 367, 390, 398, 403 

Integrated circuit 114, 298, 364 

Interlocks 226, 254, 305, 360, 377, 379 

Intermachine arcing 166, 255, 384 

Intermittent duty 153, 198-199, 222 

Interpoles, direct current motor 148 

Interrupting capacity 18, 224-229, 232-234, 238, 268, 

305, 310, 313, 325, 329, 336 

Interrupting current 225, 229, 232-234, 238, 268, 305, 

310, 313, 325, 329, 336 

Interruptors 284 

Interwinding faults 298, 311, 378 

Intrinsic safety 382, 383, 391, 393, 394-395 

Inverse-time characteristic 230, 235, 243-244, 276, 328 

Inverters 351-352 

Ionization by collision 227, 233-234 

IR-discharge voltage 293, 337 

Isolation transformers 342, 344 



-J- 



Jackets, cable 183, 185-187, 193-195, 199, 203, 

205-206, 208, 210, 393 
Jogging 90, 141 



Junction 23-24, 36, 38-39, 43, 45, 78, 104-107, 109- 

114, 164, 194, 226, 289, 346, 349, 359 
Junctions, semiconductor 112 



428 



-K- 



Page 

Kelvin double bridge 123, 398 

Kinds of protection 248, 256, 312, 362 



Kirchhoffs current law 
KirchhofFs voltage law 



Page 



23-25, 32, 35, 38-40, 79, 100, 110 

22-24, 36-38, 42, 47, 57, 67, 76, 146, 149 



-L- 



Lagging phase angle 47, 146 

Laminations, core 

rotor 130 

stator 132, 137-139 

transformer 311 

Lead-acid batteries 368-370, 374-376, 381 

Lead entrances, enclosures 387-389, 391-392 

Leading phase angle 47 

Leakage current 190, 206-207, 356, 360, 375, 399 

Lenz's law 27, 136 

Let-go current 161-162, 181 

Let-through energy 236 

Lightning 159, 163, 166, 175, 256, 280-281, 290, 292, 296, 

298-301, 306, 332, 337, 339-342, 344, 379 

Lightning protection 166, 218, 301, 337 

Line-end failures 295 

Liquid-immersed transformers 262, 335, 337-338 

Liquids, insulating 307, 337, 345 

Load-break switch 226, 235, 239, 256, 305-306, 326- 

327, 329, 331, 336, 342-343 
Load centers (see Power centers) 

Load factor 103, 154, 197-199 

Loading machine 11-12 



Locked-rotor torque 139-140, 354 

Locomotives, trolley 164, 287, 367 

Long-time delay 230-232, 275 

Longwall mining 3, 11-13, 229 

Loop 5, 14, 22-23, 36-40, 43, 55, 57, 64-65, 67, 70, 

79, 102, 117, 129, 132, 137, 151, 227, 251- 
253, 317-318, 330, 349, 352, 375, 379 

Loop equations 36-38, 40, 43, 57, 64, 67, 102 

Losses 2-3, 6, 8, 66, 79, 108, 140, 149, 206, 284, 289, 308 

core 69-70, 72-73, 137, 310 

eddy-current 70, 136-137, 311 

friction 137, 146 

hysteresis 70, 285, 311 



I^R 



137 



transformer 68, 285 

windage 146 

Low-resistance groundbed 177-178, 339, 341 

Low voltage ... 3, 5, 9-10, 15, 135, 154, 165-166, 185, 190-192, 194, 

207, 216, 226-228, 232, 235-239, 244, 254-256, 

259, 261, 263, 266, 274-275, 278, 287, 

298, 305, 311, 317, 325, 342, 344, 

348, 362, 378-379, 400, 403 

Low-voltage systems 154, 181, 261, 276, 362 



-M- 



Magnetic blowout 227 

Magnetic circuit breaker 231 

Magnetic circuits 227, 274, 311, 324 

Magnetic fields 26-27, 66, 70, 77, 115, 126-127, 129- 

130, 133, 136-139, 146, 148, 156- 
157, 227-228, 231, 258, 324 

Magnetic-flux density 26-27, 66, 69, 77, 109, 129, 

131-133, 148-149, 151, 280 

Magnetic poles 26, 137-138, 143 

Magnetic starters (see Starters) 

Magnetically coupled circuits 27, 66, 105 

Magnetization curves 115, 119, 371 

Magnetizing current 69-70, 72, 284-285 

Maintenance 2-3, 6, 18, 79, 128, 135, 143, 153, 158, 182, 202, 

207, 221-222, 224, 228, 232-234, 261, 285, 307, 

316, 325, 328-329, 331-332, 334-336, 344, 350, 

352, 354, 358, 361, 366, 368, 370, 375-377, 

379-383, 391-392, 396-399, 404 

Manual starters, direct current 150 

Margin of protection 307, 337-338 

Maximum experimental safe gap 386 

Maximum power transfer 40, 42-43 

Maximum trip setting 270, 274-275, 313, 325, 363 

Medium voltage 3, 16, 135, 179, 185, 190, 194, 226-227, 231- 

232, 234, 251-252, 254, 256, 261, 268, 
271-272, 274, 305, 311, 317, 344, 378 



Megohmeter 123-124, 

Mesh, ground-bed 159, 166 

Messenger wire 

Metal-clad enclosure 

Metal oxide varister 

Meter 20, 103, 115-125, 133, 136, 176, 

290, 316, 366, 378, 386, 

ammeter 51, 72, 89-90, 115-117, 119-120, 

273-274, 309, 317, 329, 344, 

d'Arsonval 115-117, 119-120 

demand 90, 122, 

electrostatic 

megohm 

moving iron 

multimeter 

ohmmeter 22, 

power-factor 

var 90, 118, 

voltmeter 51, 90, 115-117, 119- 

127-128, 316, 344, 

volt-ohm-milliammeter (VOM) 

watt 72, 90, 115, 117-120, 122. 

watt hour 90, 122, 127 

Methane ignition 

Microelectronics 

Mine explosions 



181, 206 

, 169-171 

203-204 

326,336 

298,356 

207, 272, 

390,403 

123-124, 

380,398 

', 124-125 

127, 345 
.. 115 
.. 181 
117, 119 
117, 125 
117, 206 
.. 118 
, 127-128 
120, 125, 
398,400 
.. 398 
, 127-128 

128, 136 
.. 384 
.. 114 
.. 367 



429 



Page 



Page 



Mine gases 3, 234, 367, 395 

Mine motors 135, 153, 155-156, 222, 401 

Mine power centers 1, 5, 200, 279, 302-308, 310-311, 313, 

315-317, 319, 320, 325, 331, 335, 345 

Mine Safety and Health Administration (MSHA) 395 

Mine ventilation 12, 17, 147, 334, 342, 373, 374, 375, 380 

Mines 

shaft 13, 15 

strip 4-5, 9-11, 185, 216, 219-220, 232 

underground 3-5, 11-13, 15, 19-20, 76, 81, 103, 164, 

166, 182, 187, 190, 192, 198-202, 204, 

207, 210-211, 223, 228, 234, 238, 

254-257, 259-261, 264, 275, 279, 

281, 285, 289, 296, 301-303, 

308, 313, 319, 326, 333-334, 

340, 343, 365-367, 372, 379, 

382-383, 389, 394-395 

Minimum-oil circuit breaker 233-234 

Mining cycle 12 

Mining methods, coal 

area, strip 8-9 

contour, strip 8-9, 220 

longwall 3, 11, 13, 229 

open pit 8, 11, 19, 185, 216, 219, 255, 301, 345 

room-and-pillar 11-12 

shortwall 13 

strip 4-5, 9-11, 185, 216, 219-220, 232 

Model, circuit 53, 72 

Moisture, soil 171 

Molded-case circuit breaker 226-232, 240, 247, 256, 270, 

274, 275, 278, 312-313, 319, 
324-325, 342, 362, 365 



Momentary-current rating 225, 360 

Motor 1-4, 9, 13-14, 16-17, 19, 26, 30, 76-79, 85, 92, 103- 

104, 113-114, 125, 133-158, 160, 176, 197-198, 

200, 222, 226, 231, 233-234, 236, 238, 261, 

275-279, 283, 285, 287, 290-291, 298, 303, 

308, 313-314, 318, 332, 346, 348-356, 

360, 380, 383, 385, 389, 391, 394-404 

brushes . . 90, 129, 132, 135, 142-143, 148-150, 350, 352, 366, 399 

capacitance 268 

commutating-pole 148 

compound 148-149, 151, 153, 348 

construction 135, 148 

enclosures 134 

induction 60, 62, 84, 89, 91, 122, 135-143, 152, 156, 158, 

242, 244, 260-261, 263, 265, 267-268, 270-271, 
297, 319, 348, 351-356, 363, 365, 397 

insulation 135-139 

mine 135, 153, 155-156, 222, 401 

series 2, 4, 135, 148, 151-154, 320, 348 

shunt 148-151 

single-phase 156-157 

squirrel-cage induction 135-136, 139-140, 143, 

154, 348, 351-352 
synchronous ... 4, 89, 135, 137, 143-148, 152, 260-261, 268, 297 

three-phase 3, 84, 135-139, 154, 283 

wound-rotor induction 142-143, 283, 352, 397 

Motor control (also see Speed Control and Starters) 

Motor-generator, sets 4, 129, 147, 350 

Moving-coil instrument 115-118 

Multiple restrike 283-284 

Mutual inductance 27, 64-67, 69, 70, 75, 170, 173 

Mutual-resistance effect 168-169, 172 



-N- 



National Electrical Code (NEC) 195, 271, 278, 382, 395 

National Electrical Manufacturers Association 

(NEMA) 134, 382, 395 

Natural frequency 260, 282-283, 285-286, 287 

Negative sequence 98-102, 180, 224, 255, 261, 263, 361, 398 

Negative-sequence components 100, 361 

Network . . 2, 5, 23, 40-41, 43-44, 48, 55, 101, 166, 267-268, 356, 358 

equivalent 33, 34 

four-terminal 176 

three-terminal 33, 34 

two-terminal 33 

Network protector 7 

Neutral . . 20, 77-79, 82-83, 86, 89-90, 93, 95, 98-102, 104, 108, 125, 
147-148, 159-161, 163-166, 171, 177, 179-181, 216, 224- 



226, 236, 249-250, 255-256, 258-259, 261, 264, 268, 

272, 276, 278, 287, 290-293, 298, 301, 308-309, 

312, 314-315, 318, 321, 323-324, 333, 

340, 343-344, 346, 354, 362-363 

Neutral-shift protection 256, 258-259, 323 

Nip 14, 256-257 

Node 36, 38^»2, 55, 64-65 

Node equations 38, 40 

Normal transient 282 

Norton's theorem 40, 43, 44, 57 

N-p-n transistor 109-113 

Nuisance tripping 141, 166, 184, 216, 251, 253, 270-271, 290, 

314, 318, 330, 360, 363-364, 367, 380 
Null 122-123 



-o- 



Ohm 21-22, 42, 53, 93, 116, 123, 162, 174, 361 

Ohmmeter 22, 117, 206 

Ohm's law 21, 23-25, 30-31, 35-36, 38-40, 46, 53, 261-262 

Oil circuit breaker (OCB) 5, 232, 329 

Oil-immersed transformers 294, 337 

One-line diagrams 86, 91-92, 95-96, 163, 201, 255, 262, 264, 

276-277, 302-303, 333-334, 336, 344 



Open circuit 30, 33, 43, 72, 97, 196, 206-207, 

215, 253, 289, 317, 368, 374 

Open delta 82, 120, 249, 251, 316, 329 

Open pit mining 8, 11, 19, 185, 216, 219, 255, 301, 345 

Operating mechanism, breaker 90, 228-229, 231, 233-235, 

239, 274, 286, 305, 312 



430 



Page 

Oscillograph 125, 300 

Oscilloscope 125-126 

Outby 182, 205, 377, 379 

Overburden 4, 8, 11 

Overcurrent protection 141, 232, 250, 275, 278, 356, 377, 380 

Overcurrent relay, alternating current time 90 

Overcurrent relay, direct current 90 

Overhead-line distribution 217, 219 

Overload 6, 97-98, 107, 119, 134, 141, 154, 171, 224, 226, 229- 

237, 248-249, 254-256, 258, 270-271, 273-276, 278, 

312-314, 331, 333, 335-336, 340, 354-356, 

360-362, 364, 365, 378, 380, 392-394 



Page 

Overload protection 107, 255-256, 271, 273-276, 278, 312, 

314, 331, 340, 355, 362, 364- 
365, 378, 380, 392, 393 

Overtravel, relay 278, 328 

Overvoltage ... 7, 64, 88, 90, 119, 159-161, 224, 240, 242, 246-247, 

260, 278-280, 283-287, 290, 292, 295, 298, 300, 

306-307, 311, 322, 329, 332, 336-337, 

354-356, 358, 361, 364, 379, 398 

Oxide film, aluminum 174 

Ozone resistance 185 



-P- 



Packing gland 192-193, 205, 387-388, 392 

Packing-gland lead entrance 387-388 

Parallel circuits 24-25 

Parallel-ground path 252, 379 

Parallel resonance 64 

Parallel-series circuits 143, 368 

Paralleling reactors 109 

Partial-discharge 185-186, 192, 398, 403 

Peak inverse voltage 105, 322 

Peak load 4, 103, 249, 270 

Peak voltage 132, 282, 286, 295-2%, 306 

Percent quantities 93 

Percent ratio error 242, 247, 272-274 

Permanent splice 207 

Permissible equipment 382, 387, 391-393 

approval 382, 391 

explosion-proof 382-383, 387, 391 

intrinsic safety 382, 393 

mobile equipment 387, 392 

Per-phase reduction 84-85, 90 

Per-unit quantities 93, 309 

Phase 

angle 46-47, 51, 52, 54, 59-61, 90, 92, 146, 

157, 244, 247, 249, 363, 399 

balance 83, 88 

current 79-80, 83-84, 92, 101, 108, 120, 250 

protection 248, 256 

sequence 90, 120, 122, 124, 137 

Phase-sensitive short-circuit protection 362-365 

Phase-sequence indicator 122-124 

Phasor 51-54, 56, 61, 77, 91, 97-98, 100, 146, 198, 201, 287 

Pilot wire 88, 90, 318 

interlocks 306 

monitoring 187, 251-254, 305-306, 317, 319, 378-379 

Plants, power . . 8, 17-19, 129, 334, 340, 342, 382, 389, 394-395, 397 

Plugging 141, 191, 350 

Plugs and receptacles (see Connectors) 

P-n junctions 104-107, 110, 112-114, 346, 349 

P-n-p transistor 109-110, 112-114 

Polar relay 242 

Polarity of windings 65 

Polarizing diode 257-258 

Poles, motors 

nonsalient 130 

salient 130, 136, 138, 143, 244 

Polychlorinated biphenyls 307, 329 

Polyvinylchloride (PVQ insulation 390 



Portable cables (see Cables) 

Portable substation 3, 5, 19, 166, 303, 332, 342 

Positive sequence 77, 98-100, 101-102, 255-284, 261, 263 

Positive-sequence components 100 

Potential difference 20, 23, 29, 115, 161, 171, 280, 290, 377 

Potential gradient 163, 166, 169-172, 175, 178, 287, 300, 339 

Potential relaying . . . 246-247, 249-250, 255, 278, 315, 333, 340, 344 
Potential transformers . . 75, 86, 118, 120, 126, 237, 243, 246, 248 

Pothead 87, 191-192 

Power 

apparent 59-60, 62-63, 66, 68, 74, 81, 

83, 93, 103, 118, 246, 319 

average 21, 50, 59-60, 63, 72-73, 80, 81, 83 

complex 59, 60-62, 68, 73-74, 80-81, 83, 84, 

103, 117, 119-122, 125, 197, 319 

imaginary 59-60 

reactive 59-60, 62-63, 81-82, 97, 118, 319 

real 59-60, 62 

three-phase ... 59, 76, 79-80, 81-82, 98, 103, 144, 154, 157, 281 

Power centers 1, 3-5, 8, 11-13, 16, 19-20, 33-34, 36, 44, 

48, 65, 76, 81, 94-95, 193, 199-201, 228- 
229, 279, 295-296, 310-317, 335, 344 

alternating current 255, 280, 302-307, 310, 313, 315- 

317, 319-320, 325, 331, 345 

alternating current-direct current 320, 325 

breakers 257 

bus 312, 319 

couplers 5, 191, 254, 304, 317, 379 

disconnect switch 5, 13, 226, 254, 256, 305- 

306, 326-327, 336, 338 

fuses 254-257, 278, 295, 306, 312, 314, 316, 

322, 329, 332-333, 336, 377-380 

grounding ... 1, 5, 13, 164-165, 190, 252, 255-257, 308, 311-312, 

314-315, 317, 319, 321, 325, 329, 333, 344, 379 

instruments 11, 75, 173-174, 176, 270 

surge arresters 295-296, 306-307, 310, 332, 334, 337-338 

transformers 200, 308, 311 

Power circuit breaker 75, 226-228, 232, 274-275, 329, 331 

Power factor ... 59, 61-63, 80-82, 84-85, 90-91, 117-120, 127-128, 

138-139, 146-147, 154, 197-198, 201-202, 237, 

246, 270, 282, 313, 319-320, 363, 399, 403 

correction 283, 332 

Power rectifier 90, 93, 107 

Power systems (also see Distribution) 

Power transformers 66, 70-71, 73, 75, 82, 86, 179-180, 238, 

276, 297, 307, 308, 310, 325, 337, 
340-341, 344, 377, 378-379 



431 



Page 



Page 



Power zeners 298 

Prefault voltage 262, 264 

Preparation plants 8, 17-19, 129, 334, 340, 342, 

382, 389, 394-395, 397 

Pressure relay 337 

Pressure-relief devices 388 

Prestrike transient 286, 295-296 

Primaiy-selective distribution 9 



Propagation velocity 288 

Protective relaying 75, 104, 118, 125, 240, 244-246, 248-249, 

254-256, 259, 273, 278, 312-313, 326, 
328, 331-333, 336, 340, 343, 345- 
346, 352, 356, 359, 362, 365-366 

Protective relays 90, 119, 160, 248, 259, 272, 344, 356, 365-366 

Pull-in torque 145 

Pull-out torque 146 



-Q- 



Quality factor, a 64 



Quick-break, quick-make mechanism 226, 229 



-R- 



Radial system 5-6, 10, 326, 333-334 

Rail 2-5, 12-14, 164-166, 204, 211, 215-216, 220, 

252, 256-257, 303, 320, 367, 392, 397 

Rail bond 165, 211, 215-216, 252 

Ratchetting, relay 360 

Rating 82, 86, 102-103, 105, 107, 118-119, 134, 140-141, 149, 

153-154, 179-180, 187, 207, 222, 239, 241, 254, 

260-266, 272-279, 290, 292, 305-310, 312-317, 

321-322, 325, 331, 337, 349-350, 362 

cable 159, 185, 195, 199 

circuit breaker 226-229, 231-233, 270, 292, 313-314, 329 

grounding resistor . . 179, 224, 247, 255, 272, 278, 312, 314-315 

motor 135, 142, 261 

switching apparatus 8, 224-226, 240, 255, 257, 264, 

267-268, 284, 295-296, 335 
transformer (see Transformer rating) 

voltage 71, 135, 144, 185, 200, 232, 236, 239, 

246, 262, 292-293, 295-296, 305-307, 
321-322, 329, 355-356, 379, 401 

Ratio error 246, 272-274 

Reactance 47, 54-57, 59, 63-64, 68, 71, 75, 93-94, 

96, 118, 140, 145-146, 160, 200-201, 
261-265, 308, 310, 321, 335 

capacitive 54, 297 

inductive 267, 271 

leakage 69 

of cables 265 

of conductors 139 

of motors 266, 267, 270-271 

Reactive power 59-60, 62-63, 81-82, 97, 118, 139 

Reactors . . 87, 109, 114, 252, 258-259, 310, 321-322, 324, 370, 372 

Real numbers 4&49 

Real power 59-60, 62 

Receiver 76, 90, 318 

Reciprocity 4fr42 

Recloser 335-336 

Recorders 124-126 

Recovery voltage 282, 283, 295 

Rectifiers 13, 90, 104-111, 117, 128, 164-166, 

211, 228, 232, 256, 258, 290, 
307, 310, 320-325, 364, 397 

battery charging 367, 371-373, 377, 379, 380-381 

control circuitry 75, 141, 151, 154, 316, 324, 326- 

327, 329, 330, 331, 349, 370 



full-wave 370 

bridge 143, 370 

three-phase 108, 348 

half-wave 106, 108, 329, 346-347 

mercury arc 370 

mine 3, 323 

overloads and faults 97-98, 224 

ratings 308 

silicon 108-109, 147, 370 

silicon-controlled 113, 346 

three-phase 107-108 

thyristor 237, 346, 348 

transformers 105-108, 321 

Reduced-voltage starters 141, 352, 365 

Reed relays 358 

Reels, cable 9, 187, 204, 222, 257, 367, 393 

Reference node 38-40 

Reference phasor 51, 54, 56, 60-61, 68, 73, 97 

Reflected wave 288-289, 294 

Refracted wave 288-289 

Regulation, voltage 1, 3, 7, 18, 73-74, 76, 85, 197, 

199, 200, 298, 310, 319 

Relays 14, 75, 86, 89, 104, 118-119, 125, 143-145, 150- 

151, 157, 159-161, 180-181, 224, 226, 230- 

231, 263, 274-279, 287, 290, 301, 306, 

311, 313-316, 323-324, 326, 334, 336- 

337, 340-341, 343-346, 353-354, 366, 

370, 376, 379-380, 396, 398, 402 

alternating current time overcurrent ... 90, 166, 243, 245, 247, 

249-250, 256, 270-271, 

328-330, 334, 358, 

360-362, 364-366 

clapper 241-243, 356 

contacts 358, 363 

cylinder 242, 245 

direct current ground-fault 259, 323 

direct current overcurrent 90, 247, 256, 361 

differential 244-247 

directional 88, 90, 240, 242, 244 

electromechanical 240-241 

ground protective 90, 165, 179, 248-249 

hybrid 356, 358, 360 

induction disk 136, 242-243, 249, 270-273, 

278, 328-329, 331, 361-362 



432 



Page 



Page 



instantaneous 241, 243, 255, 328, 364, 368 

overcurrent 240, 243, 248 

overtravel 278, 328 

residual 249-251 

solid-state 240, 356, 358-360, 361-362, 364-366 

tap settings 243-244, 249, 271-273, 331, 361 

terminology 240 

thermal 240-241 

time-delay 248 

undervoltage 240-248 

voltage 240, 246 

Reliability 5-6, 12-13, 18-19, 135, 153, 221, 251, 307- 

308, 312, 332, 334, 354, 356, 358, 
361-362, 365, 368, 377, 396 

Reluctance 69, 311 

Residual relaying 249-251 

Resins 135, 139, 221, 361, 390 

Resistance 2, 20-21, 23, 24, 26, 32-36, 3945, 50-57, 59, 61, 

64, 66, 83, 93, 98, 101-102, 105-108, 111-112, 

116, 132, 138, 141-143, 146, 157, 160-162, 

166, 167-169, 176, 178-181, 190, 192, 200, 

211, 213, 215-216, 224, 241, 246-249, 

252, 254-256, 260, 262, 264-270, 273- 

276, 283-298, 306-308, 310-315, 317- 

321, 338-346, 352, 356, 359-365, 

368-370, 375, 378-379, 382, 394 

alternating current 46-47 

cable 123, 164, 207 

conductor 22, 31, 70, 123, 139-140, 159, 195 

direct current 184 

ground-bed 124, 163, 170, 172, 174, 177, 300, 339 

insulation 123, 185, 207, 396, 399, 400, 402-403 

parallel 25, 30, 43, 117 

series 29-31, 117, 123, 125 

winding 68-73, 148-153 

Resistivity 166-171, 174-187, 190, 339, 394 



earth 178, 180-181, 301, 341 

material 22 

Resistor 22, 24-25, 29-30, 33, 3741, 46-48, 54, 89, 

112, 114-118, 123, 126, 141-142, 145, 

150, 160, 171, 246, 253-256, 258-260, 

292, 309, 311-318, 323, 332-334, 

338, 352, 358, 363, 370, 398 

current-limiting 165 

grounding 90, 161, 163-165, 179-180, 224, 247, 249- 

250, 253, 255, 268, 272, 278, 287, 297, 

312, 314-315, 323, 333-334, 340- 

341, 343-344, 354, 361 

starting 151 

Resonance 287, 319 

parallel 64 

series 64 

Response, frequency 361 

Restrike 98, 226, 282-285, 287 

Retentivity 70 

Reverse bias 105-106, 110, 112, 114, 322, 346 

Reversing controls, motor 150 

Rheostats 90 

Ripple voltage 108-109, 132 

Rod, ground bed 168 

Roof bolter (drill) 81, 153, 190 

Room-and-pillar mining 11-12 

Root-mean-square (rms) 50-51, 59 

Rotating lines 51 

Rotating magnetic field 136-137, 146, 157 

Rotor 130-133, 135-138, 141, 142, 156-157, 

244, 260, 353-354, 383, 391 

bars 136, 139-140, 142, 404 

construction 140 

cylindrical 143-145 

wound 142-144, 283, 352, 397 

Rules, ground bed 178 



-s- 



Sacrificial anode 178 

Safety factor, borehole cables 198, 203, 261-262, 268, 275, 

278, 328, 367, 373-374, 
380, 386-387, 393 

Safety grounding 160, 163-164, 166, 170, 178- 

181, 334, 338-341, 343, 377 

Sag, overhead line 217 

Salient pole 130, 136, 138, 143, 244 

Saturable reactor 252, 258-259, 324, 370-372 

Saturable transformer 258 

Saturation current 104-105, 110 

Saturation curves 273 

Schedule 142, 194, 207, 302, 367, 370, 381, 391-392 

Schedule 2G 259, 374 

Secondary-selective system 6, 334 

Secondary-spot network 5, 7 

Sectionalizing unit 326 

Segment 3, 13-14, 17, 20, 76, 93-94, 113, 132, 147-148, 150, 

153-154, 201, 211, 254, 276, 284, 288, 296, 320 

Selective relaying 159, 224 

Selective -system operation 18, 224 

Selector device 90 



Selenium suppressor 322 

Self-inductance 27, 65-66 

Semiconductor devices 104, 114, 356 

Semiconductor fuses 237 

Semiconductor shields 210 

Sensitive earth-linkage system 362 

Sensors 207, 221, 354, 356-357 

Series circuits 22-25, 54, 64, 76, 94, 124 

Series motors 4, 151 

Service factor 134 

Settings, maximum instantaneous 275 

Shaft mines 13, 15 

Shell, coupler 319 

Shield 13, 98, 183, 189, 194-196, 200-201, 206, 210, 

224, 231, 281, 292, 299, 325, 342, 363, 383 

conductor 186-187 

electrostatic 311 

Faraday 298, 311, 379 

insulation 186-187, 190 

nonmetallic 186 

overhead lines 298 



433 



Page 



Page 



transformer winding 67, 74-75, 95, 119, 246, 

308-310, 321, 336, 379 

Shielded cable 9, 11, 186-187, 189-191, 199, 

204, 210, 219, 260, 319 

Shielding failure 281 

Shock, electric 160, 161-162, 180-182, 186, 222, 316- 

317, 339, 366, 367, 377, 380 

Short circuit . . 23, 33, 40, 42-44, 57, 73, 90, 95, 97-98, 107, 146, 180, 

184-185, 197, 207, 224, 228-229, 232-236, 239, 

248-249, 254-256, 261, 265, 270, 274-278, 

286, 289, 298, 306, 309-312, 328, 333, 

336-339, 361-364, 367, 389-390, 395 

Short-circuit currents 44, 73, 202, 225, 229, 232, 236-237, 

260-264, 270, 275, 277, 310, 313, 
321, 328, 331, 336, 344 

calculations 278 

Short-circuit protection 231, 236-237, 249, 256, 270-271, 273- 

276, 279, 312, 314, 319, 331, 340, 

353, 356, 362-363, 365-366, 

378, 380, 392 

Short-time ratings 336 

Shorting switch 317 

Shovels 4-5, 8, 10, 85, 147, 158, 166, 182, 

190, 270-271, 281, 351-352 

Shunt 4, 42, 90, 106, 116-118, 125-126, 132, 148-151, 231- 

234, 240, 242, 247-248, 256, 292, 296, 298, 306, 
313, 315, 324, 348, 362, 371, 376, 398-399 

Shunt field 132, 150-151 

Shunt motor 148-151 

Shunt trip 231-234, 248, 298, 306, 313, 315, 362 

Shuttle car 2-3, 12, 15, 29-30, 81, 96-97, 151, 153, 165, 183, 

185, 187, 190, 196, 198, 204-206, 208-209, 
222, 308, 320, 350-351, 367, 391 

Silicon-controlled rectifier (SCR) (thyristor) 113, 346 

Silicon diode 3, 108, 165, 287, 370 

Silver-sand fuse 238 

Single-cage rotor 140 

Single-ended substation 333 

Single-line diagram 86, 256 

Single-phase analysis 84 

Single-phase motors 156-157 

Slip . . . 129, 136-140, 144, 153-154, 156, 209-210, 348, 352, 387-388 

Slip rings 90, 129, 131, 135-136, 142-143, 257, 352 

Snubber 298, 323, 356 

Soil heating 170-171, 180 

Soil resistivity 166-169, 171, 174-175 

treatment 177 

Solenoids 27, 141, 231, 234, 241-242, 312-313, 356 

Solid-material fuse 238 

Solid-state belt starters 319, 352-353 

Solid-state device . . 104, 107, 125, 232, 237, 346, 356-360, 376, 379 

Solid-state motor control 356 

Solid-state trip elements 312-313, 365 

Solidly grounded system 276 

Source transformation 40, 42-45, 57 

Source transformer 249, 278, 343-344, 363 

Sources 19, 23-24, 39-43, 45, 48, 55, 69, 76-79, 84, 97, 

100, 110, 159-160, 185, 225, 242, 247, 260- 
263, 270, 280, 295, 302, 343, 348, 383, 394 

Spacing, overhead lines 217 

Span, overhead lines 217-221 

Spark gap 98, 292, 298, 363 

Sparkover voltage 292-296, 306-307, 337-338 

Specifications 2, 93, 135, 158, 185-186, 194-195, 197-198, 200, 

204, 208, 211-212, 229, 246, 280, 302-303, 
307, 314, 335, 370, 381, 383, 387-391 



Speed control, motor 

chopper 349-350 

compound-motor 148-149, 151, 153, 348 

induction motor 60, 62, 84, 89, 91, 122, 135-143, 152, 

156, 158, 242, 244, 260-261, 263, 

265, 267-268, 270-271, 297, 319, 

348, 351-356, 363, 365, 397 

series-motor 2, 4, 135, 148, 151-154, 320, 348 

shunt-motor 148-151 

solid-state 356 

Ward-Leonard system 152-153, 350-351 

wound-rotor motor 142-143, 283, 352 

Speed-torque curves 133 

Splices, cable 186, 195, 207-210 

Squirrel-cage motor 139-140, 154 

single-phase 135, 348 

three-phase 135-136, 143, 351 

Squirrel-cage rotor 145, 156, 404 

Standard burdens, current transformers 119 

Starters, alternating current 16, 113, 125, 145, 150, 238-239, 

298, 351, 356, 366, 402 

across-the-line 141, 149 

autotransformer 74, 114, 141, 159 

belt conveyor 319, 353-355, 360, 365 

magnetic 87, 142 

reduced voltage 141, 352, 365 

solid-state 319, 352-353 

synchronous motor 4, 89, 135, 137, 143-148, 

152, 260-261, 268, 297 

wound rotor 142-143, 283, 352, 397 

Starting torque 77, 133, 135, 140, 144-145, 149, 151, 157, 354 

Static device 358-362 

Static lines 341-342 

Station arrester 293-294 

Stator 130-132, 135-146, 156-157, 244, 260 

Steady-state response 48 

Storage battery 42, 367-368, 373, 381 

Stranding conductor 185, 404 

Stray ground current 367 

Stress cone 191, 194 

Strip-chart recorders 124-126 

Strip heaters 329 

Strip mining 4-5, 9-11, 185, 216, 219-220, 232 

Stuffing box 387-388 

Substation ... 1, 3, 11, 13, 15, 17, 19, 76, 83, 91-92, 96, 179, 191, 

199-201, 217, 220, 233, 237-238, 256, 265, 281, 

293-294, 299, 302-303, 319, 326, 337, 343 

area 339-341, 344-345 

bus 255 

capacity 7-8, 334 

circuit breakers 10, 333 

connections 328 

design of 332, 342 

double-ended 334 

economics 8 

fuses 333 

grounding 9, 166, 180-181, 249, 338, 342, 345 

loads 335, 342 

location 8, 344 

portable 5 

relaying 249, 336 

switches 5, 336 

transformers 128, 163, 307, 331, 334-336, 342 



434 



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Page 



unit 307 

Substitution 34, 40, 43 

Subtransient reactance 261-268, 270-271 

Subtransmission 8-9, 13, 342 

Superposition 40-41 

Suppressor, transient 298, 322 

Surface leakage, battery 375-379 

Surface mining 2, 4-5, 7-11, 13, 19, 104, 159, 166, 182-183, 

185, 190-191, 194, 199, 204, 206-207, 210- 

211, 216, 218-220, 237, 255, 260-261, 

270, 275, 279, 281, 289, 296, 303, 

319, 326, 343-344, 382, 394 

Surface oxide film, aluminum 192 

Surge arrester 22, 87, 163, 292-2%, 298-299, 301, 306- 

307, 310, 327, 332, 334, 337-338, 3% 

Surge capacitor 260, 268, 282, 295-297, 307, 332 

Surge impedance 166, 285, 288-289, 295, 299 

Susceptance 47, 55, 64 

Switch 9-10, 87, 89-90, 105, 116, 120, 121, 144, 147, 151, 

157, 191, 215-216, 224, 231, 235, 239-240, 253, 

257, 280-281, 285, 291, 316-319, 329, 331- 

332, 342-344, 348-349, 370, 376-380 

disconnect 5, 13, 220, 226, 237, 254, 256, 

305-306, 326-327, 336, 338 

interlock 150, 305-306, 378-379 

interrupter 226 

load-break 226 

push-button 330, 356, 360, 392-393 

ratings 260 

Switches, fusible 235, 237, 254, 306 

Switchgear 14, 159, 261, 267, 307 

Switchhouses ... 1, 5, 10, 13, 15, 128, 166, 182-183, 193, 201, 219, 
224, 255-256, 263, 296, 302, 305, 335, 345 

breakers 9, 326, 329 

connections 285, 295 

couplers 191 

design of 331 

relaying 278, 326, 328, 334 

Switching apparatus 8, 224-226, 240, 255, 257, 264, 

267-268, 284, 295-296, 335 
Switching skid 326 



Switching transients 234, 281-282, 286-287, 295 

Switchyard 8-9 

Symbols 21-24, 27-28, 30, 38, 47-48, 55, 66, 

74, 82, 86-88, 98-99, 104-105, 

109, 112-115, 240, 343, 346 

Symmetrical components . . 98-103, 179, 250, 261-262, 266, 268, 404 

Symmetrical currents 263 

Synchronous generators 131, 260 

Synchronous motors 4, 89, 135, 137, 143-148, 

152, 260-261, 268, 297 

Synchronous reactance 261, 265 

Synchronous speed 137, 139-140, 145-146, 153, 156-157 

System 1-22, 30, 33-34, 36, 48, 50, 62-63, 65, 71-73, 75-86, 

90-103, 105, 118-120, 129, 131, 135-136, 139-140, 

147, 152-154, 157-173, 175, 178-182, 186-188, 

190-192, 194-195, 197, 199-201, 203-206, 210- 

211, 216-266, 275-290, 292-302, 304-308, 

310-313, 316-318, 321-326, 329, 331-345, 

355-359, 362-367, 370, 372, 375-377, 

379-383, 386, 394-398, 400-401, 404 

alternating current-direct current 164 

control 4, 349, 352-353, 354 

diode-grounded 164, 187, 257-258 

expanded radial 5 

grounding 160, 165, 378 

high-voltage 160, 165, 239, 251, 253, 268, 271, 293, 298 

load 98, 215, 319 

low-voltage 154, 181, 261, 276, 362 

medium-voltage 179, 227, 231, 268, 271-272 

power (see Power systems) 

primary-loop 6 

primary-selective 6, 9 

reliability 5, 356 

resistance-grounded ... 179, 276, 296, 317, 321, 340, 362, 365 

secondary-selective 6, 334 

simple radial 9 

solidly grounded 276 

trolley 5, 14, 154, 164-166, 182, 211, 215, 

228, 256-258, 270, 323-324, 350 
ungrounded 160, 272, 276, 287, 293 



-T- 



Tachometer 353-354 

Tap changers 333-334, 337 

Tap settings, relay 243-244, 249, 271-273, 331, 361 

Tape recorders 125-126 

Temperature 68, 90, 105-107, 110-112, 134-135, 153, 155-156, 

174, 177, 180, 183-184, 194, 196-199, 201, 206- 

208, 217, 222, 226, 230, 240-241, 246, 271, 

274, 307, 310, 312, 349, 354, 358-359, 

369, 370, 374-375, 383-384, 386-387, 

393-395, 398-400, 402, 404 

coefficient of resistance 22, 175 

effect on ampacity 195 

ground bed 171 

limits 235 

permissible enclosure 389 



ratings 185, 195 

Temperature rise 68, 134-135, 153, 155-156, 171, 195, 222, 

230, 235, 271, 307, 310, 312, 383 

Temporary splice 207, 393 

Testing 123-124, 222, 246, 314, 329-331, 360-362, 

375, 387-388, 391, 395, 399-403 

cable 206-207 

system 316 

transformer 316 

Thermal conductivity 171 

Thermal device 90, 241 

Thermal-magnetic breaker 229, 313 

Thermal overload 271, 337, 354, 356 

Thermionic emission 226, 234 

Thevenin's theorem 43-44, 57 



435 



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Page 



Three-phase 3, 57, 59, 76-85, 90-93, 97-103, 107-108, 120-121, 

124-125, 131, 135-141, 143-144, 154, 156-157, 

164, 179, 216, 225, 229, 239, 246, 248, 250, 

255-256, 260-261, 263-264, 266, 268, 270, 

274-275, 278, 281, 283-284, 294-296, 

306, 308-309, 316, 319-320, 322, 

328-329, 335-336, 346, 348, 

350-352, 363, 370, 372, 377 

Three-phase generators 77, 131, 135 

Three-phase rectifier 107-108 

Three-phase systems 76-81, 83, 85, 99, 131, 136, 

239, 246, 294, 296 

Thumper 207 

Thyristor 113-114, 237, 287, 298, 323, 346-360, 370-372 

Time-current curves 229, 236, 271, 360 

Time-delay relay 242, 248-249, 255-256, 263, 266, 278 

Time dial setting 244, 270, 278 

Time-domain reflectometer (TDR) 207 

Time rating 140, 179, 312 

Tinned conductors 200 

Tolerance 194, 197, 199, 331, 360, 374, 399 

breaker settings 270, 274-275, 278-279 

voltage 227-228, 232, 248 

Torque 77, 115, 117-118, 122, 124, 133-135, 137-138, 

139-157, 242, 244, 348, 352-354, 361 
Torque-speed characteristics 

direct current motors 147 

inductor motors ... 60, 62, 84, 89, 91, 122, 135-145, 152, 156, 

158, 242, 244, 260-261, 263-265, 267- 

268, 270-271, 297, 319, 348, 

351-356, 363, 365, 397 

synchronous motors 4, 89, 135, 137, 143-148, 

152, 260-261, 268, 297 

wound-rotor motors 142-144, 283, 352, 397 

Touch-and-step potentials 171, 338-339 

Track ... 8, 154-155, 164, 166, 211, 215-216, 220, 303, 367, 376, 390 

grounding 165 

resistance 192 

Track bonds 165, 216, 367 

Traction locomotives 154, 155, 156 

Trailing cables (see Cables, portable) 

Transducers 115, 119, 126, 224, 240 

Transfer functions 349 

Transformation . . 3, 13, 33-34, 40, 42^15, 55, 57, 83, 92, 164, 332 

Transformers ... 27, 59, 64, 78, 81-83, 85-87, 89-90, 94-96, 99, 103, 

111, 129, 139, 158-159, 163-164, 166, 183, 200, 

224, 228, 237-238, 240-241, 243-251, 254-256, 

259, 264, 265, 271-273, 276-278, 293-298, 

312-316, 318-328, 331-338, 340-342, 350- 

354, 361, 370-372, 377-381, 402-404 

auto 74, 141, 159 

air-cooled 134 

basic impulse insulation level (BIL) 290, 366 

burden 119-120, 328 

capacitance 296 

capacity 68, 81, 103, 295, 306, 316, 334, 356 

class 307 

construction 130, 311 

control 306, 315-316, 324, 326, 329, 343 

cooling 334 

core 70 

core loss 69-70, 72, 73, 137 

current . . 83, 89, 108, 118, 243-244, 246, 249, 272-273, 328-329 



delta-delta 83, 92, 261, 308, 343-344 

delta-wye 83, 92, 261, 308, 343-344 

design 69, 173, 310 

distribution 5, 7, 13, 75, 262, 297, 303, 309, 398 

dry-type 2, 93, 301, 337, 379 

eddy-current loss 70, 136-137, 311 

efficiency 146 

electrostatic shielding 311 

equivalent circuits ... 40, 57, 68, 70-71, 73, 76, 85, 94-95, 113- 

114, 260-261, 266, 284-285, 291 

exciting current 69-73, 94-95, 274, 295, 308 

Faraday shielding 298, 311, 379 

fault current 310 

ferroresonant 371-372, 381 

forced-cooled 334 

frequency 72-73 

grounding 79, 179-180, 308-309, 321, 340, 343-344 

high-voltage 305, 319, 334 

ideal 66-71, 73-74, 91, 108, 363 

impedance 94 

input impedance 66-67, 110, 112, 125 

inrush current . . . 150, 244, 270, 276, 285-286, 295-296, 306, 337 

instrument 118-119, 270 

insulation 159 

interwinding faults 298, 311, 378 

isolation 342, 344 

kilovoltampere rating 261, 307-308, 310 

leakage reactance 68-69, 71, 75, 310 

liquid-immersed 262, 335, 337-338 

magnetizing current 69-70, 72, 284-285 

oil 294, 337 

open-delta 82, 120, 249, 251, 316, 329 

performance 68 

potential 75, 86, 118, 120, 126, 237, 243, 246, 248 

power 200, 308, 311 

protection 118, 271, 301, 334, 336-337, 344 

pulse 354 

ratings 246, 308-309 

rectifier 105-108, 321 

regulating 301, 345 

resistance loss 118 

saturable 258 

saturation 108 

single-phase 67, 82-83, 180, 316, 335 

source 249, 278, 343-344, 363 

substation 128, 163, 307, 331, 334-336, 342 

taps 308 

temperature 310 

three-phase 82, 92, 179, 295, 308-309, 336, 350, 377 

three-winding 95 

transient voltages 280, 282, 283-287, 291-292, 

295-296, 298, 301 
turns ratio 66-68, 70, 72, 75, 83, 91-92, 95, 196, 262, 

270, 272-273, 291, 314, 317, 329 

unit substation 307 

utilization factor 107 

voltage 74, 91-92, 247, 251, 262, 295, 305, 

308-309, 313, 316, 321, 363, 370 

windings 67, 74-75, 95, 119, 2%, 

308-310, 321, 336, 379 

wye-delta 85, 90, 92, 179 

wye-wye 90-92, 308 

zig-zag 79, 89, 179-180, 308-309, 321 



436 



Page 



Page 



Transient 20, 48, 125, 159-160, 186, 192, 197, 224, 234, 252, 

261, 265, 278, 298, 300-301, 306-307, 311, 322, 
329, 336-337, 350, 354-356, 360, 379, 395 

characteristics 282 

protection 292, 295-296, 332 

sources 280 

Transient reactance 261 

Transient response 48 

Transient suppressors 298 

Transistors 109-114, 298, 356-359 

Transmission 2-3, 7-8, 65, 76, 127, 162, 181, 216, 218, 222- 

223, 259, 282, 287, 293, 301, 384, 390, 395 

Traveling waves 286-288, 289-290, 294, 296, 342 

Triac 356-360 



Trip-free relay 90, 229 

Trip settings 270, 274-275, 303, 313, 325, 363 

Tripping, circuit breaker ... 75, 184, 229, 231, 235, 240, 248, 251, 

253-254, 255, 298, 305, 315, 318- 

319, 323-324, 330, 336, 354, 

356, 359, 367 

Tripping, element . . . 228-231, 240, 248, 274, 310, 312-313, 362, 365 

Trolley rectifiers 232, 322, 325 

Trolley systems 5, 14, 154, 164-166, 182, 211, 215, 

228, 256-258, 270, 323-324, 350 

Trolley wire 2, 13, 164, 211, 214-215, 256, 270, 278 

insulators 166, 367 

Two-layer earth structures 176, 181 



-u- 



Unbalanced loads 10, 79-80, 82, 97 

Unbalanced system 100-102 

Underground coal mining ... 4-5, 15, 19-20, 76, 81, 103, 164, 166, 

187, 190, 192, 198-202, 204, 207, 

210-211, 223, 228, 234, 238, 

254-257, 259, 261, 264, 275, 

279, 281, 285, 289, 296, 

301-302, 313, 319, 326, 

333-334, 340, 343, 365- 

366, 372, 379, 382-383, 

389, 394, 395, 397 

advance versus retreat 182, 303 

continuous 3, 12-13, 16, 201, 308 

conventional 11-12, 367 

longwall 3, 11-13, 229 

room-and-pillar 11-13 



shortwall 13 

Undervoltage release 231, 256, 306, 312, 316, 318, 330, 342 

Underwriter's Laboratories 228, 259, 382, 387, 389, 394-395 

Undergrounded systems 160, 272, 276, 287, 293 

Unidirectional thyristor control 348, 354 

Uniform design 331 

Unit substation 5, 10-11, 200, 249, 307, 332, 341-342 

Units 3, 5, 14, 17, 20-22, 27, 46, 50, 55, 60, 82, 103, 107, 114, 

116, 133, 166, 182, 218, 220-221, 226, 228-230, 232- 

233, 236-237, 241, 243, 246-249, 256, 274-275, 292, 

295, 302-303, 307-308, 314, 318, 326, 328, 335 

Unloaded power factor 399, 403 

Utility, electric system 7 

Utilization 4-6, 9-10, 14, 96, 107, 153-154, 158, 180, 182, 185, 

197, 199, 224, 263, 277, 291, 293, 295, 303, 307- 
308, 310-313, 316-317, 319-320, 355, 362, 366 



-V- 



Vacuum circuit breaker (VCB) 



Var 



232, 234-235, 284-285, 295- 

296, 329, 332, 343 

90 



Varmeter 90, 118, 127-128 

Velocity propagation 288 

Ventilation 129, 381-382, 395, 397 

battery boxes 373-374, 375-376, 378, 392 

charging stations 373, 375, 377, 380 

mine 12, 17, 147, 334, 342, 373-375, 380 

Visible disconnects 326 

Volt 5, 20, 23, 25-26, 36, 40, 42, 73, 90, 93, 95, 279 

Voltage . . . 2-7, 9-11, 13, 17-48, 50-57, 59-61, 63-68, 70-86, 90-102, 

104-112, 114-122, 124-132, 134-137, 139, 141, 145- 

146, 149, 150, 152-154, 156, 159-162, 165-166, 

171, 175, 178-181, 185-187, 190-192, 194-195, 

197, 199-202, 207, 210-211, 216-217, 220- 

222, 224-229, 231-232, 234-236, 238-241, 

243-244, 246-248, 250-258, 260-265, 268, 

270, 272-275, 277, 280-301, 303, 305- 

317, 319-325, 329-330, 332, 335-339, 

342-344, 346, 348-361, 363-365, 368- 

373, 378-380, 395, 398-401, 403404 



arc 227, 286, 306, 324, 338 

cable ratings 159, 185, 195, 199 

control 346, 349 

drop calculation 199 

drop maximum 199 

gradients 175, 178, 180 

rating 71, 135, 185, 194, 200, 232, 236, 239, 

246, 262, 292-293, 295-2%, 305-307, 
321, 322, 329, 355, 356, 379, 401 

regulation 3, 7, 18, 19, 73, 74, 76, 88, 197, 

199, 200, 298, 310, 319 

relay 240, 246 

ripple 108-109, 132 

standard 216, 246 

three-phase 77 

transformer ratings 246, 265, 308-309 

Voltage classes 216 

Voltage gradient 175, 178, 180, 280, 339 

Voltampere 60, 66, 73-74, 93, 107, 119, 247, 274 

Voltmeter 51, 90, 115-117, 119-120, 125, 

127, 128, 316, 344, 398, 400 
Vulcanization 184-185, 203, 207, 210 



437 



-w- 



Page 



Page 



Ward-Leonard system 152-153, 350-351 

Watt 60, 106, 118, 137 

Watthourmeter 90, 122, 127-128, 136 

Wattmeter 72, 90, 115, 117-120, 122, 127-128 

Wave sloping 295-296 

Waves, traveling 286-290, 294, 296, 342 

Wenner array 176, 178, 181 

Wheatstone bridge 122 

Windage 137, 146 

Windings 26-27, 86, 94, 108-109, 118, 152-153, 156- 

157, 159-160, 180, 240-241, 249-250, 258, 

271-272, 298, 317, 322, 335-337, 341, 

352, 354, 363, 370-371, 377- 

378, 383, 398, 400, 402 



armature 130-132, 136, 146, 148-149, 151 

compensating 148 

response to transients 290-291 

transformer ... 67, 74-75, 95, 119, 246, 308-310, 321, 336, 379 

Windows and lenses 388 

Wire (see Conductor) 

Wooden poles 216-218, 281 

Wound-rotor motor 142-143, 283, 352, 397 

Wye connections ... 34, 77, 79-87, 93, 100, 138, 141, 179, 249, 258, 

261, 296, 308-309, 321, 329, 336, 343, 363 

Wye-delta transformations . . . 33-36, 55, 77-87, 90, 92-93, 100, 108, 

120, 138, 141, 179-180, 249, 261, 
308-309, 321, 336, 344, 363 



-Y- 



YBUS load-flow analysis 268 



-Z- 



ZBUS fault analysis 268 

Zener diodes 105, 298 

Zero-sequence relaying 249-251, 255-257, 272, 314-315, 

329, 340, 343, 362, 365 



Zig-zag transformers 79, 89, 179-180, 308-309, 321 

Zones of protection 254, 259 



* U.S. GOVERNMENT PRINTING OFFICE : 1991 - 287-709 



